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Structural evaluation of impact damaged prestressed concrete I girders repaired with Fiber Reinforced Polymer (FRP) materials

University of Florida Institutional Repository

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STRUCTURAL EVALUATION OF IMPACT DAMAGED PRESTRESSED CONCRETE I GIRDERS REPAIRED WITH FIBER REINFORCED POLYMER (FRP) MATERIALS By KIMBERLY ANN LAMMERT A THESIS PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF ENGINEERING UNIVERSITY OF FLORIDA 2003

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Copyright 2003 by Kimberly Lammert

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ACKNOWLEDGMENTS Completion of this thesis and the research associated with it would not have been successful without the help and guidance of a number of individuals. The author would like to thank the chair and members of her committee Dr. Perry S. Green, Dr. H.R.Hamilton, and Dr. Andrew J. Boyd for their guidance and assistance. Others deserving of thanks due to contributions of resources and knowledge are Dr. Thomas Sputo, Tony Michael, Jeff Brown, Johnny Fung, Jennifer Burtz, Vanessa Grillo, Koren Britt, and Viswanath Urala. The author would also like to thank the Florida Department of Transportation for their financial assistance. Lastly, the author would like to thank all of her close friends and family. Their enduring support proved to be of the utmost value. iii

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TABLE OF CONTENTS Page ACKNOWLEDGMENTS.................................................................................................iii LIST OF TABLES............................................................................................................vii LIST OF FIGURES.........................................................................................................viii ABSTRACT.....................................................................................................................xiii CHAPTER 1 INTRODUCTION........................................................................................................1 Objective of Current Study...........................................................................................3 Relevance of Current Study..........................................................................................4 2 PREVIOUS RESEARCH.............................................................................................9 Fiber Reinforced Polymer Properties and Behavior.....................................................9 Condition Assessments...............................................................................................13 Repairs with Fiber Reinforced Polymer Systems.......................................................16 3 DESCRIPTION OF EXPERIMENTAL STUDY......................................................49 Description of Test Specimens in Damaged State......................................................49 Test Setup for Four Point Bending Tests....................................................................50 Repair Designs and Design Properties.......................................................................51 Repair Application Procedures...................................................................................52 4 EXPERIMENTAL TEST RESULTS.........................................................................70 Test 1 Undamaged Control Results............................................................................71 Test 2 Damaged Control Results................................................................................73 Test 3 RJWatson Repair Results.................................................................................75 Test 4 Air Logistics Repair Results............................................................................78 Test 5 UF Sprayed Repair Results..............................................................................81 Test 6 Edge Composites Repair Results.....................................................................83 Evaluation of Experimental Observations..................................................................83 Effect of Fiber Reinforced Polymers on Stiffness...............................................83 Effect of Fiber Reinforced Polymers on Capacity..............................................85 iv

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Effect of Fiber Reinforced Polymers on Strains on Bottom Centerline of Test Specimens........................................................................................................87 Other Effects Initial Imperfections and Bond..................................................87 Summary of Test Specimen Failure Modes........................................................88 5 ANALYTICAL TEST RESULTS............................................................................110 Calculation of Effective Prestress.............................................................................110 Theoretical Cracking Capacity for Test Specimens in Undamaged and Damaged State without Repair.............................................................................................114 Theoretical Ultimate Capacity for Test Specimens in Undamaged and Damaged State without Repair.............................................................................................115 Theoretical Ultimate Capacities of Test Specimens 3, 4, 5, and 6...........................118 Support Conditions...................................................................................................120 Theoretical Midspan Deflection at the Theoretical Ultimate Capacity of the Girder in the Undamaged State............................................................................121 Comparison of Theoretical and Experimental Capacities and Deflections..............122 Computer Program Implementation for all Test Specimens....................................125 6 DESIGN RECOMMENDATIONS AND PRELIMINARY DESIGN SPECIFICATIONS...................................................................................................133 7 SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS FOR FUTURE RESEARCH.............................................................................................................138 Summary...................................................................................................................139 Conclusions...............................................................................................................139 Future Testing Recommendations for Measurements to Assess Adequacy and Performance of FRP Repair.................................................................................140 APPENDIX A REPAIR DESIGNS..................................................................................................142 RJWatson Repair Design..........................................................................................142 Air Logistics Repair Design.....................................................................................143 UF Repair Design-As Designed...............................................................................154 UF Repair Design-As Built......................................................................................158 B THEORETICAL EVALUATION TOOLS..............................................................163 Effective prestress.....................................................................................................163 Cracking Capacity for Test Specimens in Undamaged and Damaged State............165 Ultimate Capacity for Test Specimens in Undamaged and Damaged State.............167 Ultimate Capacity for Repaired Test Specimens 3, 4, 5, and 6................................170 Support Conditions...................................................................................................174 Deflection at Ultimate for Test Specimen 1.............................................................175 v

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REFERENCES................................................................................................................176 BIOGRAPHICAL SKETCH...........................................................................................180 vi

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LIST OF TABLES Table page 1-1: Summary of Bridges in Florida Damaged due to Vehicular Impact............................5 2-1: Qualitative Comparison of Carbon, Aramid and Glass Fibers...................................29 2-2: Comparison of Load, Deflection, Stiffness, and Crack Width...................................29 2-3: Comparison of Midspan Deflection Before and After FRP Application...................30 2-4: Comparison of Reinforcing Bar Stresses Before and After FRP Application...........30 2-5: Comparison of Initial Stiffness and Strength for Various Retrofits...........................31 2-6: Comparison of Stiffness Before and After Retrofit....................................................31 2-7: Comparison of Ductility Ratios..................................................................................31 3-1: Comparison of Design Repairs for Test Specimens 3, 4, 5, and 6.............................57 4-1: Experimental Capacities, Deflections, Strains, and Initial Flexural Stiffness of All Test Specimens at Cracking Load............................................................................90 4-2: Experimental Capacities, Deflections, and Strains of All Test Specimens at Maximum Load........................................................................................................90 5-1: Comparison of Experimental and Theoretical Capacities of Test Specimens 1, 2, 3, 4, 5,and 6 at Cracking.............................................................................................128 5-2: Comparison of Experimental and Theoretical Capacities of Test Specimens 1, 2, 3, 4, 5 and 6 at Maximum Load.................................................................................128 6-1:Environmental Reduction Factor for Various FRP Systems and Exposure Conditions .............................................................................................................137 vii

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LIST OF FIGURES Figure page 1-1: Number of Bridges in NBI versus Vertical Clearance of Selected 79000 Bridges......6 1-2: Number of Bridges in NBI versus Vertical Clearance of Selected 79000 Bridges by Functional Classification............................................................................................6 1-3: Chaffee Road Bridge Overall Damage Looking East, Westbound Traffic..................7 1-4: Close-up of Damage on the West Fascia Girder in the Vicinity of the Diaphragm.....7 1-5: Close-up of Damage on the East Fascia Girder Showing Severed Prestressing Strands........................................................................................................................8 2-1: Stress-Strain Diagrams for Mild Steel and Carbon, Glass, and Aramid Fiber Reinforced Polymers................................................................................................32 2-2: FRP Placement for Increase in Shear Capacity..........................................................32 2-3: Flexural Failure Modes for FRP Strengthened Beams...............................................33 2-4: Shear Failure Modes for FRP Strengthened Beams...................................................34 2-5: Load versus Midspan Deflection for Specimen 1 with 54 ksi yield steel..................34 2-6: Load versus Midspan Deflection for Specimen 2 with 54 ksi yield steel..................35 2-7: Load versus Midspan Deflection for Specimen 3 with 45 ksi yield steel..................35 2-8: Load versus Midspan Deflection for Specimen 4 with 54 ksi yield steel..................36 2-9: Load versus Midspan Deflection for Specimen 5 with 45 ksi yield steel..................36 2-10: Load versus Midspan Deflection for Specimen 6 with 45 ksi yield steel................37 2-11: Moment Deflection Relationship for Six Specimens...............................................38 2-12: Moment versus CFRP Strain for Repaired Specimens.............................................39 2-13: Moment versus Reinforcing Bar Strains for Six Specimens....................................40 viii

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2-14: Moment versus Concrete Compressive Strain for Six Specimens...........................41 2-15: Overall Damage to I-680 over CR L34, near Beebeetown, Iowa............................42 2-16: Moment Deflection Relationships for Beams 1 and 2..............................................42 2-17: Experimental and Theoretical Load Deflection Relationship for Beams 1 and 2 ...43 2-18: Load Deflection Relationship of Fully-Wrapped Beams.........................................43 2-19: Relationship Between Number of FRP Layers to Moment Capacity at Steel Yield and Ultimate...................................................................................................44 2-20: Load Deflection Relationships for Fully-Wrapped and Partially-Wrapped Sections with 3 Layers of FRP.................................................................................44 2-21: Theoretical and Experimental Midspan Deflection for the Analyzed Sections.......45 2-22: Load Deflection Relationship for Beams 1, 2, 3, and 4...........................................46 2-23: Damage to Bridge A10062, St. Louis County, Missouri.........................................46 2-24: Overall Damage to Bridge A4845, Jackson County Missouri.................................47 2-25: Close-up Damage to Bridge A4845, Jackson County Missouri...............................47 2-26: Load Midspan Deflection Relationship for the Control, Repaired with Fabric, and Repaired with Spray Specimens........................................................................48 3-1: Typical Profile and Cross-Sections of Test Specimens in Undamaged State............58 3-2: Typical Cross-Section Details for Test Specimens 2, 3, 4, 5, and 6..........................59 3-3: Typical Damage for Test Specimens 2, 3, 4, 5, and 6 Showing Removed Concrete and Two Severed Prestressing Strands.....................................................................59 3-4: Stress-Strain Curves for Removed Strands from Test Specimen 2............................60 3-5: Load, Shear, and Moment Diagrams for All Four-Point Bending Tests....................60 3-6: Typical Instrumentation for Four-Point Bending Tests..............................................61 3-7: Test Specimen 3 Repair Design.................................................................................63 3-8: Test Specimen 4 Repair Design.................................................................................64 3-9: Spray Equipment: Gun (G), Catalyst (C), and Resin (R)...........................................65 ix

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3-10: Detailed View of Spray Gun: Resin (R), Catalyst (C), Spray (S), and Chopper (Ch)...........................................................................................................................65 3-11: Spray Technique Procedure Showing How the Glass Fibers, Resin, and Catalyst for Resin are Combined in Mid-Air.........................................................................66 3-12: Test Specimen 5 Repair Design...............................................................................67 3-13: FRP Sampling Cores for Test Specimen 5...............................................................68 3-14: Test Specimen 6 Repair Design...............................................................................69 4-1: Test Specimen 1 Load versus Midspan Deflection Curve.........................................91 4-2: Test Specimen 1 Measured Deflection Profile...........................................................91 4-3: Test Specimen 1 Strain Profile at North Load Point..................................................92 4-4: Test Specimen 1 Strain Profile at Midspan................................................................92 4-5: Test Specimen 1 Strain Profile at South Load Point..................................................93 4-6: Test Specimen 1 Load versus Measured Strain Along Bottom Centerline of Beam.........................................................................................................................93 4-7: Test Specimen 2 Load versus Midspan Deflection Curve.........................................94 4-8: Test Specimen 2 Measured Deflection Profile...........................................................94 4-9: Test Specimen 2 Strain Profile at North Load Point..................................................95 4-10: Test Specimen 2 Strain Profile at Midspan..............................................................95 4-11: Test Specimen 2 Strain Profile at South Load Point................................................96 4-12: Test Specimen 2 Load versus Measured Strain Along Bottom Centerline of Beam.........................................................................................................................96 4-13: Test Specimen 3 Load versus Time Showing Loading and Unloading Cycles.......97 4-14: Test Specimen 3 Load versus Midspan Deflection Curve.......................................97 4-15: Test Specimen 3 Measured Deflection Profile.........................................................98 4-16: Test Specimen 3 Strain Profile at North Load Point................................................98 4-17: Test Specimen 3 Strain Profile at Midspan..............................................................99 4-18: Test Specimen 3 Strain Profile at South Load Point................................................99 x

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4-19: Test Specimen 3 Load versus Measured Strain Along Bottom Centerline of Beam.......................................................................................................................100 4-20: Test Specimen 3 During Loading Showing Crack Propagation on South End of Girder at the Bottom Layer of Prestressing............................................................100 4-21: Test Specimen 3 at Failure Showing Concrete Cover Separation at South End....101 4-22: Test Specimen 4 Load versus Midspan Deflection Curve.....................................101 4-23: Test Specimen 4 Measured Deflection Profile.......................................................102 4-24: Test Specimen 4 Strain Profile at North Load Point..............................................102 4-25: Test Specimen 4 Strain Profile at Midspan............................................................103 4-26: Test Specimen 4 Strain Profile at South Load Point..............................................103 4-27: Test Specimen 4 Load versus Measured Strain Along Bottom Centerline of Beam.......................................................................................................................104 4-28: Test Specimen 4 at Failure Showing Separation of FRP.......................................104 4-29: Test Specimen 4 at Failure Showing Shearing of FRP Stirrup..............................105 4-30: Test Specimen 5 Load versus Midspan Deflection Curve.....................................105 4-31: Test Specimen 5 Measured Deflection Profile.......................................................106 4-32: Test Specimen 5 Strain Profile at North Load Point..............................................106 4-33: Test Specimen 5 Strain Profile at Midspan............................................................107 4-34: Test Specimen 5 Strain Profile at South Load Point..............................................107 4-35: Test Specimen 5 Load versus Measured Strain Along Bottom Centerline of Beam.......................................................................................................................108 4-36: Test Specimen 5 at Failure Showing Rupture of FRP at Midspan.........................108 4-37: Test Specimen 5 at Failure Showing Rupture of FRP at Midspan on Tensile Face of Girder.........................................................................................................109 5-1: Geometric Properties of the Cross-Section a) Girder without the Slab; b) Slab; and c) Girder with the Slab....................................................................................130 5-2: Steel Areas and Location for a) an Undamaged Section; and b) a Damaged Section....................................................................................................................131 xi

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5-3: Normalized Moment versus Deflection for Test Specimens 1, 2, 3, 4, and 5..........132 5-4: Moment Curvature Comparison of Theoretical Data from Computer Program and Experimental Data..................................................................................................132 xii

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Abstract of Thesis Presented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Master of Engineering STRUCTURAL EVALUATION OF IMPACT DAMAGED PRESTRESSED CONCRETE I GIRDERS REPAIRED WITH FRP MATERIALS By Kimberly Ann Lammert December 2003 Chair: Perry S. Green Major Department: Civil and Coastal Engineering In Florida, any vehicle over 13 feet 6 inches is defined as an over-height vehicle and requires a permit to travel on any road system in Florida. Although regulations are present and enforced that does not ensure that collisions with bridge structures due to over-height vehicles will not occur. Impact damage due to over-height vehicles can lead to structure collapse, reinforcement damage, girder misalignment, steel yielding, connection failure, reinforcement exposure, concrete spalling, and concrete cracking. After a collision has occurred in which an emergency repair is required, a fast, inexpensive, effective, and easy repair is preferred. The impetus for this research is to assist the Florida Department of Transportation (FDOT) in establishing a quality products list (QPL) of acceptable methods for the repair of impact damaged prestressed concrete bridge girders using fiber reinforced polymer (FRP) materials. The necessity of a quality products list was realized after the Chaffee Road Bridge over Interstate 10 in Jacksonville was severely damaged on July 6, 2001, xiii

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from an over-height vehicle and needed immediate repair. At that time, the FDOT had only one company or method to repair the bridge, and because most FRP methods are proprietary to a company, realized the necessity of needing other approved alternatives. The existence of a quality products list will enable qualified companies to be certified by the FDOT to perform emergency repairs to vehicular damaged bridge girders using FRP materials. Full-scale girder tests of six type II AASHTO girders were performed. The six tests represented an undamaged control specimen, a control specimen with simulated damage, and the four specimens with simulated damage that were then repaired with different FRP systems. The FRP systems varied in material types, laminate properties, application procedures, reinforcing schemes, and repair design procedures and assumptions. The specimens were tested to failure to determine moment and shear capacities, as well as deformation and ductility behavior of the undamaged, damaged, and repaired girders. Through experimental and analytical comparisons of the behavior of the repaired girders to the undamaged girder, the specific FRP systems were evaluated for acceptance to the FDOT Quality Products List. The FRP systems were evaluated structurally based on their ability to restore the undamaged shear and moment capacity of the prestressed concrete girder, type of failure mode, cost of repair system including installation, and ease of installation. xiv

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CHAPTER 1 INTRODUCTION The 1990 edition of the American Association of State Highway and Transportation Officials A Policy on Geometric Design of Highways and Streets states that the minimum vertical clearance for highways is recommended as 14.5 feet and is desired as 16.5 feet (AASHTO 1990). Not all overpass bridges currently meet these standards either due to being constructed before these standards were issued or paving overlays of the underlying roadway (Bridge Engineering Software and Technology [BEST] 2001). Figure 1-1, assembled by the National Bridge Inventory (NBI), shows the clearance heights for 79000 bridges that have service both over and under the structure in the U.S.. Figure 1-2 shows the clearance heights of these 79000 bridges in terms of their functional classification either interstates and freeways or other arterials, collectors, and locals. In Florida, any vehicle over 13 feet 6 inches is defined as an over-height vehicle and requires a permit to travel on any road system in the state (Florida Department of Transportation [FDOT] 1998). By obtaining a permit the driver of an over-height vehicle is provided knowledge of how to get a the desired destination on roads with sufficient vertical clearance, and the FDOT is provided information about who, where, and when the road system will have an over-height vehicle in case of bridge damage. Low clearance postings are placed in advance of every bridge or structure with a minimum vertical clearance of 14 feet 6 inches or less (FDOT 1999 revised 2002). In addition, low 1

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2 clearance postings are placed on the structure for every bridge or structure with a minimum clearance of 13 feet 6 inches or less. Although regulations exist and are enforced does not ensure that collisions with bridge structures due to over-height vehicles will not occur. Impact damage due to over-height vehicles can lead to structure collapse, reinforcement damage, girder misalignment, steel yielding, connection failure, reinforcement exposure, concrete spalling, and concrete cracking. After a collision has occurred in which an emergency repair is required, a fast, inexpensive, effective, and easy repair is preferred. Table 1-1 lists the bridges in Florida that have been damaged due to vehicular impacts over the past two years. Table 1-1 provides for each of the impacted bridges the approximate date of impact, district number, intersecting roads, and if any other impacts due to over-height vehicles have been reported. Many research projects have been conducted to determine what criteria should be used for the assessment of damaged concrete structures (Shanafelt and Horn 1980, Shanafelt and Horn 1985, Zobel et al. 1996, Arockiasamy and Barbosa 2000). Establishment of these assessments would provide the necessary means to ensure that if the structure is to be repaired that all potential problems are investigated, and a proper and effective repair could be installed. Numerous experimental studies of laboratory repairs and in field repairs with fiber reinforced polymer (FRP) systems on concrete structures have been conducted (Sen and Liby 1994, Arockiasamy 1995, Tedesco et al. 1998, Klaiber et al. 1999, Mayo et al. 1999, Shahawy and Beitelman 1999, Labossieere et al. 2000, Tumialan et al. 2001, Scheibel et al. 2001, Boyd and Banthia 2001, Spadea et al. 2001). All of the experimental research

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3 has shown a promising future for fiber reinforced polymer systems for the repair of damaged or deteriorated concrete or steel structures. Objective of Current Study The impetus for this research is to assist the Florida Department of Transportation in establishing a quality products list (QPL) of acceptable methods for the repair of impact damaged prestressed concrete bridge girders using fiber reinforced polymer (FRP) materials. The necessity of a quality products list was realized after the Chaffee Road Bridge over Interstate 10 in Jacksonville was severely damaged on July 6, 2001 from two over-height vehicles and needed immediate repair. The Chaffee Road Bridge, built in 1960, is comprised of five Type III AASHTO girders with a 7 deck and an original underclearance of 15-6. Figure 1-3 shows the overall damage done to the Chaffee Road Bridges prestressed concrete girders including entrance and exit damage in the background and foreground of the picture, respectively. Figures 1-4 and 1-5 show more detailed views of the damage where the vehicles exited from under the bridge overpass and where the vehicle made first contact with the bridge overpass and the resulting damage. These figures show significant concrete cover loss to all girders, cracks extending into the web for all girders, and severing of prestressing strands to the exterior girders. At that time, the FDOT had only one company or method to repair the bridge, and being that most FRP methods are proprietary to a company, realized the necessity of needing other approved alternatives. The existence of a quality products list will enable qualified companies to be certified by the FDOT to perform emergency repairs to vehicular damaged bridge girders using FRP materials.

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4 Relevance of Current Study Full-scale girder tests of six type II AASHTO girders were performed. The six tests represented an undamaged control specimen, a control specimen with simulated damage, and four specimens with simulated damage that were then repaired with different fiber reinforced polymer systems. The FRP systems varied in material type, laminate properties, application procedures, reinforcing schemes, and repair design procedures and assumptions. The specimens were tested to failure to determine moment and shear capacities, as well as deformation and ductility behavior of the undamaged, damaged, and repaired girders. Through experimental and analytical comparisons of the behavior of the repaired girders to the undamaged girder, the specific fiber reinforced polymer systems were evaluated for acceptance to be placed on the FDOT Quality Products List. The FRP systems were evaluated structurally on their ability to restore the undamaged shear and moment capacity of the prestressed concrete girder, the type of failure mode that occurred, the cost of repair system including installation and ease of installation.

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5 Table 1-1: Summary of Bridges in Florida Damaged due to Vehicular Impact DateDistrictFacility CarriesIntersectsMultiple hits1/15/20022I-95SR 206YES7/6/20012I-10CR 115C (Chaffee Road)YES2/16/20014I-95Linton Blvd (SW 12th St)NO4/24/20024I-95SR 708 (Blue Heron Blvd)YES8/13/20017SR 618 (Crosstown EXPY)34th StYES9/5/20028SR 91 TPKSR 826 Palmetto EXPYNO5/29/20018SR 91 TPKSR 91 Jupiter InterchangeNO

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6 Figure 1-1: Number of Bridges in NBI versus Vertical Clearance of Selected 79000 Bridges (BEST 2001) Figure 1-2: Number of Bridges in NBI versus Vertical Clearance of Selected 79000 Bridges by Functional Classification (BEST 2001)

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7 Figure 1-4 Figure 1 5 Figure 1-3: Chaffee Road Bridge Overall Damage Looking East, Westbound Traffic Figure 1-4: Close-up of Damage on the West Fascia Girder in the Vicinity of the Diaphragm

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8 Figure 1-5: Close-up of Damage on the East Fascia Girder Showing Severed Prestressing Strands

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CHAPTER 2 PREVIOUS RESEARCH Fiber Reinforced Polymer Properties and Behavior Previous research has shown a promising future for the use of fiber reinforced polymer systems as an efficient means to strengthen or retrofit concrete structures in order to resist increased design loads or repair damage (American Concrete Institute [ACI] 2002). Fiber reinforced polymer systems are lightweight, easy to install, noncorrossive, and inexpensive, when compared to prior repair techniques such as bonded steel plates, external post-tensioning, or concrete and steel jacketing (International Federation for Structural Concrete [FIB] 2001). When determining whether to repair or replace a structure, it is crucial that if failure of the FRP system occurs that it does not result in failure of the structure. A fiber reinforced polymer system consists of fibers and a polymer matrix, namely a stress-bearing component, and a stress-transferring component (FIB 2001). Fibers are typically made of carbon, glass, or aramid and all exhibit an almost linear elastic behavior to failure (ACI 2002). Figure 2-1 compares the stress-strain curves for carbon, glass, and aramid fiber reinforced polymer systems to that of mild steel. The high tensile strength of an FRP system provides increases in axial strength by fully wrapping the cross-section, in shear capacity by placing the fibers perpendicular to potential shear crack development as shown in Figure 2-2, and in flexural capacity by placing the fibers in the tension areas. FRP systems can be used not only to increase flexural strength but can also be used to 9

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10 increase the ductility in columns by providing confinement through wrapping the entire cross-section of the column. Polymer matrices or adhesives are available in epoxies, polyesters, and vinyl-esters. Each type of fiber and adhesive has its advantages and disadvantages, which are dependent on application procedure, type of structure, reason for rehabilitation, and the configuration of the system. Table 2-1 provides a qualitative comparison of different fiber composites. An effective fiber reinforced polymer system consists of fibers and an adhesive working together so that the fibers can take on load from the original structure through an adhesive bond. If the bond is not capable of transferring the required load, the necessary capacity of the fibers can not be developed. Fibers can be manufactured in a continuous or discontinuous form (FIB 2001). Also, the fiber strands can be manufactured into unidirectional, bi-directional, or multidirectional patterns or orientations forming a fabric which is capable of resisting and distributing load in the manufactured directions (ACI 2002). A fabric is the arrangement of fibers bound together in two or more directions. Fiber reinforced polymer systems are divided into their different application techniques and can be classified as a wet-lay-up system, pre-impregnated system, pre-cured system or a sprayed system (ACI 2002, FIB 2001). A wet-lay-up system consists of applying a saturating adhesive in-place to a dry reinforcement forming a cured laminate on a structural member. A pre-impregnated system consists of an uncured fabric that has been previously impregnated off-site with a saturating resin which is then bonded to the surface of a structural member with or without an adhesive depending on system requirements. A pre-cured system is applied using an adhesive to bond the off-site

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11 previously cured and formed laminate to a structural member. The spray technique combines the fiber, resin, and catalyst for the resin at the time of application and is sprayed directly onto the structural member (Boyd and Banthia 2001). The placement of the system on a structure and the orientation of the fibers determines the type and amount of enhanced capacity which the rehabilitated structure will be capable of resisting (ACI 2002, FIB 2001). To resist moment in a flexural member, the fibers or fabric should be placed in the tension areas of the structural member and oriented in the direction of the highest tensile stresses. To resist shear in a flexural member, the fibers or fabric should be applied to the sides of the member and oriented either perpendicular to latent or visible shear cracks or at 45 from the longitudinal axis of the member as shown in Figure 2-2. When increasing the moment capacity of a structural member it is necessary to check that sufficient shear capacity is present to resist the corresponding increase in shear loads or additional shear reinforcement must be provided. There are many possible failure modes for beams flexurally strengthened with FRP systems. If the ends of the FRP system are sufficiently anchored to the tension side of a flexural member, brittle failure will occur when the beam reaches its ultimate flexural capacity through tensile rupture of the FRP or concrete crushing as shown in Figures 2-3a and 2-3b, respectively (Teng et al. 2002). Figure 2-3c shows that a brittle failure in shear can occur if the flexural capacity exceeds the shear capacity of the strengthened member. Brittle and premature failure can result from debonding due to high interfacial stresses at or near the ends of an FRP system as shown in Figures 2-3d and 2-3e, respectively. Brittle and premature failure can also result from debonding due to high interfacial

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12 stresses from a flexural or flexural-shear crack formed away from the ends of a FRP system as shown in Figures 2-3f and 2-3g, respectively. In addition, other failure modes are possible such as failure of the adhesive, slip at the concrete substrate to adhesive interface and slip at the adhesive to fabric interface or by a combination of any of the modes described above. If the retrofitted member is a prestressed element, additional failure modes can occur such as a prestressing strand steel failure. Possible failure modes for shear strengthened beams are shear tension with FRP rupture, and shear tension without FRP rupture where the FRP debonds as shown in Figures 2-4a and 2-4b, respectively (Teng et al. 2001). Both failure modes for shear strengthened members are brittle that can lead to an abrupt and catastrophic failure. Equations, examples, and methods for determining FRP contribution have been developed and included in the ACI (2002) code for reinforced concrete but not for prestressed concrete. The only guidelines are that strain compatibility regarding the state of strain in the prestressed member should be used to calculate FRP capacity increases and that rupture of prestressing strands should be considered as a failure mode (ACI 2002). The absence of code standards means that design repairs are being performed under increased engineering intuition, which could lead to incorrect assumptions. Guidelines for detailing have been developed to avoid bond-related failures. Suggested details incorporate the use of mechanical anchorages to assist in stress transfer, tapering FRP lengths for multiple ply laminates, and providing sufficient overlap for splicing FRP plies (ACI 2002).

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13 Condition Assessments Shanafelt and Horn (1980) reported on the findings of an extensive research and compilation of statistics of all cooperating state departments of transportation. The responding departments of transportation showed that of the 23,344 prestressed concrete bridges in those states, an average of 201 bridges were damaged each year. The most significant finding was that over 80 percent of the damage to prestressed concrete bridges was due to over-height vehicles. The investigation produced the necessity for a standard method to evaluate damaged bridges, classify the amount and types of damage, and how to make decisions on the repairs. The types of repairs investigated in the report were repairs with externally bonded reinforcing bars, external post-tensioning using strands or bars, mild steel external sleeves, and internal strand splices. These repairs were investigated through analytical calculations on beams classified as severely damaged. Shanafelt and Horn (1985) further investigated the findings of the research by performing repairs in the laboratory with external post-tensioning, mild steel external sleeves, and internal strand splices. The purpose of the experiments was to establish standards for evaluating damage to bridges, determining the amount of damage, and designing and applying appropriate repair techniques. Although some of the repair techniques investigated are no longer used solely for repairs, internal strand splicing is still currently being investigated as part of a potential repair technique system. Zobel, Jirsa, Fowler, and Carrasquillo (1996, 1997 revised) reported on field inspections and laboratory experiments of an impact damaged prestressed concrete girder. The damaged fascia girder was removed from the bridge and used to evaluate non-destructive tests on impact damaged girders, and evaluate repairs with internal strand splicing and concrete patching. Inspection of the girder showed cracking extending into

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14 the web with undamaged exposed prestressing strands. The non-destructive techniques used to evaluate the amount and type of damage to the girder were the rebound hammer, impact echo, and spectral analysis of surface waves. In addition to the above assessments, the girder was load tested without any repairs up to service load. The results showed that the girder had a lower flexural stiffness and had larger deflections at locations of significant damage and higher strains in strands with cross-sections that had significant damage. The girder was repaired using several cast-in-place repair methods: latex-modified vertical overhead mortars, single-component latex-modified mortars, and fiber reinforced silica fume modified mortars over different damaged sections. Epoxy was then injected to all remaining cracks and voids. After the repairs were completed, the girder was reevaluated with the rebound hammer which showed a sound repair and with the impact echo and spectral analysis of surface waves which showed that structural integrity was generally restored. The beam was then load tested again up to its service load. The results indicated that the girder had a higher flexural strength though it still had larger deflections at locations with significant damage. The slope of the load versus tendon strain curve showed significant improvement after the repair was completed. The authors also investigated the effectiveness of internal strand splicing. Four strands were intentionally severed at midspan to be spliced with four different splicing assemblies. To determine the effectiveness of the splice, the girder was taken through a series of load tests in which all four strands had been severed, all four strands spliced, two splicing assemblies removed, and with all four splices removed. The load tests demonstrated that the strain in the undamaged strands decreased when the damaged

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15 strands were spliced suggesting a redistribution of strain to the spliced strands and the capability of the splices to restore service load capacity. Arockiasamy and Barbosa (2000) assessed techniques for the condition evaluation and repair of concrete bridges. The research provides condition evaluation, selection of repair types, repair design procedures, repair construction procedures, and case studies of implemented repairs for bridge deck, substructure and superstructure elements. The effective types of repairs that were considered in this study for prestressed girders were external post-tensioning, internal splicing of strands, metal sleeve splices, and replacement. The Bridge Engineering Software and Technology (BEST) Center (2001) reported on a study in which the purpose was to determine the extent of over-height vehicle collisions and identify ways to prevent over-height vehicle collisions. A survey was sent to every state requesting the following: average clearance height of bridges, maximum vehicle height, fines for over-height vehicles, clearance postings, and collision data. Only 29 states replied to the survey. Of the responding states only 12 provided statistics on over-height collisions. Only 17 states stated that they keep records of over-height vehicle collisions. The researchers determined from the submitted data that the number of over-height vehicles increased by 1.3% per year from 1995 to 1999. In Maryland, 309 of the 1496 susceptible bridges or 20 percent have been damaged from over-height vehicles. Of the 309 bridges that have been damaged 106 or 34% cross over an interstate highway. To increase bridge clearances seven states or 24 percent indicated that they had ground down the pavement that resulted from repeated resurfacing. The state of Georgia

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16 responded that a program was ongoing to raise all existing interstate bridges to a minimum clearance height of 16-6. Repairs with Fiber Reinforced Polymer Systems Although a significant amount of research has been conducted on fiber reinforced polymer (FRP) strengthening systems of reinforced concrete elements, little research has been conducted on the use of FRP for the repair of prestressed concrete elements. This section contains the last 10 years of research concerning repairs with FRP systems. At present, standards for the design, application, and configuration of fiber reinforced polymer strengthening systems on prestressed concrete structures have not been developed. Sen and Liby (1994) reported on the feasibility of using CFRP laminates to strengthen steel composite beams through experimental and analytical procedures. The composite beams were made of W8x24 steel sections with a 28 by 4 concrete slab. The specimens were subjected to two loading cycles in four-point bending. The first cycle consisted of loading the specimen to yield of the tension flange to simulate distress. The second cycle consisted of loading the specimen to failure after the specimen was repaired with CFRP laminates along the tension flange. Clamps were installed at the ends of the laminates to resist peeling stresses. The clamps were designed based on the stresses determined from a finite element analysis in the laminate and adhesive. The test specimens consisted of two different steel yield strengths of 45 ksi and 54 ksi and were repaired with two different laminate thicknesses of 2mm and 5mm. Specimens 1 and 2 were 54 ksi yield steel repaired with 2mm thick CFRP laminates, Specimen 3 was 45 ksi yield steel repaired with 5 mm thick CFRP laminates, Specimen 4 was 54 ksi yield steel repaired with 5 mm thick CFRP laminates, and Specimens 5 and 6 were 45 ksi yield steel

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17 repaired with 2 mm thick CFRP laminates. After testing Specimens 1, 2, and 3, additional bolts were used to further resist the stresses due to shear in the adhesive. These additional three rows of bolts were spaced 16 apart at each end of the CFRP laminate. Figures 2-5, 2-6, 2-7, 2-8, 2-9, and 2-10 show the load deflection relationships for Specimens 1, 2, 3, 4, 5, and 6, respectively. The experimental results showed that the CFRP laminates significantly increased the ultimate capacity of the steel composite sections. A higher capacity with a more ductile failure mode could also be reached by properly anchoring the CFRP laminates with the additional bolts as compared to the unanchored repaired specimens. Arockiasamy (1995) studied the flexural behavior of rectangular reinforced concrete beams with a varying number of CFRP plates on the tension face through experimental tests and analytical procedures. Six concrete specimens were tested to failure, one was a control with no FRP, one was strengthened with one layer of FRP, two were strengthened with two layers of FRP, and two were strengthened with three layers of FRP. The beam dimensions were 8 wide by 1-0 deep and were 8-0 long with a span to depth ratio of four. The beams were all tested to failure in two-point loading, and were instrumented with seven strain gauges on the bottom centerline of the beam, four strain gauges on one side of the beam, and five LVDTs equally spaced along the length of the beam. Table 2-2 compares the ultimate capacity, deflection at midspan, initial stiffness, and midspan crack width for all of the specimens, with the last number in the beam name indicating the number of layers of CFRP. Figure 2-11 shows the moment deflection relationship for all of the specimens. All of the beams with FRP systems failed due to concrete crushing at smaller deflections than the control specimen. The bonding

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18 of the CFRP plates to the tension face increased the flexural capacity, increased the flexural stiffness, reduced crack widths, and reduced the CFRP, rebar, and concrete compressive strains at midspan as the number of FRP plates increased. Moment versus strain relationships for the CFRP, steel reinforcing bar, and concrete are shown in Figures 2-12, 2-13, and 2-14 respectively. Tedesco, Stallings, and EL-Mihilmy (1998) investigated the effects of rehabilitating a deteriorated reinforced concrete bridge in the field with external bonding of FRP plates. The reinforced T-beam bridge was located on State Highway 110 near Union Springs, Alabama. The simple span bridge had significant flexural and shear cracking before repair. All of the reinforcing laminates used for repairs consisted of unidirectional fibers oriented parallel to the longitudinal axis of the plate. Three of the four beams were retrofitted with carbon fiber reinforced polymers on the tension face to increase flexural capacity and with glass fiber reinforced polymers on the sides to increase shear capacity. The remaining beam was only retrofitted with carbon fiber reinforced polymers on the tension face. Load tests were performed before and after the retrofit was applied to determine what affects the application of the FRP strengthening system had on the bridge. Retrofitting the bridge resulted in smaller strains in the longitudinal reinforcement and lower midspan deflections as shown in Tables 2-3 and 2-4. The beam with only FRP on the tension face exhibited the least reduction in strain and deflection indicating the significant effect of the GFRP on the overall structural behavior. To verify the effects that FRP laminates have on strengthening the bridge, a finite element analysis was performed. Results from the finite element analysis showed close

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19 correlation to the in-field load tests for reinforcement strains and midspan deflections with an average percentage difference of 5.7% and 5%, respectively. Klaiber, Wipf, Russo, Paradis, and Mateega (1999) reported on the repair of an impact damaged prestressed concrete bridge in Iowa in which field and laboratory tests were conducted. The impacted bridge carries Interstate 680 over County Road L34 located near Beebeetown, Iowa. The bridge consists of two separate structures of the same design and construction; one carries traffic eastbound and the other carries traffic westbound. Each bridge consists of a concrete deck set on eleven beams with varying vertical clearances. The westbound bridge was damaged while the eastbound bridge was undamaged, thus providing a means to compare in-place load tests on the damaged and undamaged structures. The three northernmost girders of the westbound bridge showed significant damage at their midspans ranging from exposed and slack multiple prestressing strands, a pre-existing severed strand, web cracking, and spalling of concrete. Figure 2-15 shows some of the damage to the bridge. As part of the study, the eastbound and westbound bridges were static load tested to determine whether they had a different response to the applied load and load distribution. Although the two northernmost damaged girders did not require replacement, they were subsequently removed from the bridge for laboratory testing of fiber reinforced polymer systems and the third northernmost girder was repaired insitu. Static load tests were again performed on the westbound bridge after the removed girders were replaced. The field test results showed that the behavior of the eastbound and westbound bridges were now basically the same.

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20 Beam 1 was tested to failure as the control specimen. Beam 1 had two severed strands, which caused the beam to fail in combined shear and slab crushing at a moment of 2067 kip-ft and corresponding deflection of 8.62. Beam 2 was damaged by severing strands to be equivalent to Beam 1. Beam 2 was then repaired with three longitudinal carbon fiber reinforced plates with additional fiber stirrups along the length to restore the loss of capacity due to the severing of two strands and loaded to failure. Beam 2 ultimately failed due to debonding of the carbon fiber reinforcing system at a moment and corresponding deflection of 2480 kip-ft and approximately 6.4. Figure 2-16 provides a comparison of the moment deflection relationships for Beams 1 and 2, the series B1W is for Beam 1 and all of the other series are for Beam 2. By taking the experimental strain in the composite at ultimate and multiplying by the area of FRP material, the researchers determined that the force in the FRP was equivalent to three and a half 250 ksi strands with a diameter of 0.5. Only two strands had been severed, therefore the design goal of restoring the lost tensile capacity of the damaged strands was reached. A 12% increase in ultimate capacity of the repaired girder (Beam 2) over the control specimen (Beam 1) and a 10% increase in ultimate capacity over the theoretical ultimate capacity of an undamaged member was reported. Mayo, Nanni, Watkins, Barker, and Boothby (1999) reported on the strengthening of a simple span reinforced concrete solid slab bridge in Iron County, Missouri. The bridge, located on a mining truck operation route, Route 32, was strengthened in order to remove weight restrictions. Through a comparison of design trucks in use at the time of design to the present and including the current condition of the bridge such as corrosion damage and cracking, it was determined that a 19% increase in flexural strength would be

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21 required to remove the weight limit posting. To ensure the effectiveness of the FRP strengthening system, two full-scale reinforced concrete beams were constructed and tested to failure. The beams were constructed to simulate the existing bridge conditions and geometry. Beam 1 was a control beam with no FRP strengthening, and Beam 2 was strengthened with a one-ply carbon fiber reinforced polymer system to verify that a 20% increase in capacity could be restored. Beam 1 failed due to concrete crushing, and Beam 2 failed due to FRP rupture. Figure 2-17 provides a comparison of the experimental and theoretical load deflection relationships for Beams 1 (before strengthening) and 2 (after strengthening). The laboratory test results showed that a 27% increase in flexural strength with a finer crack pattern was achieved over the control beam test result. The durability of the bonding of the FRP system was also investigated through static and dynamic tests, confirming that peeling of the FRP should not be an issue. The bridge was instrumented with deflection gauges and then load tested while before and after FRP strengthening. The in-field load tests indicated a slight increase in stiffness due to decreased deflection of the rehabilitated bridge. Shahawy and Beitelman (1999) studied the static performance of reinforced concrete beams strengthened with CFRP sheets. Eight T-beams of approximately 19-0 long were tested to failure in four-point bending. One specimen was tested as a control with no FRP reinforcement. Five specimens had FRP applied to the entire stem, meaning the bottom and sides of the beam, with varying layers of FRP. For the five fully-wrapped specimens, there was one specimen with one layer of FRP, one specimen with three layers of FRP, one specimen with four layers of FRP, and two specimens with two layers of FRP. The remaining two specimens had two layers FRP applied only to the bottom

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22 surface of the stem. All of the specimens were repaired with unidirectional carbon fiber sheets. The T-beams were instrumented with LVDTs at the load points, supports, and midspan and with strain gauges at the load points and midspan. The moment versus midspan deflection curve for the fully wrapped beams and control beam show an increase in capacity and a reduction in deflection of all of the repaired beams when compared to the control specimen shown in Figure 2-18. Figure 2-19 shows how the number of layers of FRP affects the flexural capacity at steel yield and ultimate for the fully wrapped beams. The partially wrapped specimens showed lower moment capacity and deflection than the fully wrapped sections presented in Figure 2-20. The partially wrapped beams failed prematurely due to concrete cover separation, demonstrating the significance of providing sufficient anchorage through wrapping the beams. A two-dimensional finite element program, which takes into account the non-linear response of the component materials was implemented. Four cross-sections were analyzed with this program representing the control specimen, the partially wrapped beam with two layers of FRP, and the remaining were fully wrapped cross-sections with two and three layers of FRP. The program resulted in very consistent values for midspan deflection for all cross-sections with slightly higher deflections and capacities at ultimate as shown in Figure 2-21. Labossieere, Neale, Rochette, Demers, Lamothe, Lapierre, and Desgagne (2000) reported on the strengthening of the Sainte-Emelie Bridge that needed upgrading in order to withstand heavy loads due to a nearby timber distributor. The Sainte-Emelie Bridge carries Route 131 over the Riviere Noire near Sainte-Emelie-de-l-Energie in Quebec, Canada. The bridge is a one-span monolithic deck and T-section reinforced concrete

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23 structure with no significant signs of deterioration such as concrete spalling or corrosion. The weight of standardized design trucks has increased since the design of this bridge. Therefore, to bring the bridge up to current standard design loads, the bridge required a 35% increase in moment and 20% increase in shear capacity. To ensure the strengthening procedure would be sufficient, numerous laboratory tests were performed on one-third scale test specimens and analytical models were evaluated. The laboratory tests included different potential climatic effects, behaviors of FRP materials, assessment of an assortment of strengthening configurations for the T-section, and the bonding capacity of the FRP system to concrete. Four beams were constructed and tested to failure. Beam 1 was a control beam with no additional reinforcement. Beam 2 was reinforced with six layers of carbon fiber for flexural strengthening on the full length with glass fiber stirrups at varying spacing with an additional continuous strip of glass at the top of the stirrups for added anchorage. Beam 3 was reinforced with one layer of carbon fiber for flexural strengthening on the full length with the glass fiber stirrup spacing of Beam 2 with an additional continuous strip of glass at the top of the stirrups for added anchorage wider than Beam 2. Beam 4 was reinforced with six layers of carbon fiber for flexural strengthening, three extending the full length and three apporximatley 3-0 shorter on both ends with glass fiber stirrups at varying spacing and an additional continuous strip of glass at the top of the stirrups for added anchorage wider than Beam 2. The control beam (Beam 1) failed in shear. Beam 2 failed in shear due to debonding of the FRP stirrups but with a 40% load capacity increase over Beam 1. Beam 3 failed when the flexural strengthening failed in tension with a 60% load capacity

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24 increase over Beam 1. Beam 4 failed when the flexural strengthening ruptured in tension with a 50% load capacity increase over Beam 1. Figure 2-22 provides the load deflection relationships for Beams 1, 2, 3, and 4. A larger deflection at failure was observed for the beams with FRP than the control beam consistent with the analytical models. The laboratory experiments showed the effectiveness of the FRP material on strengthening reinforced concrete structures. The final retrofit design for the Sainte-Emelie Bridge included three layers of carbon fiber oriented longitudinally for the full length of the beams. In addition, glass fiber stirrups with fibers oriented vertically anchored at the top with an additional glass fiber oriented longitudinally were applied to the beams. To determine the in-field effectiveness of the FRP system, the bridge was instrumented with strain gauges primarily at midspan and located on the reinforcing steel inside the beams, on the FRP strips. A displacement sensor was also placed under each beam at midspan. The bridge was load tested before and after the FRP application to observe the bridge behavior. Decreases in the midspan strains and deflections were observed after the FRP application as hypothesized in the design and demonstrated in the laboratory. Tumialan, Huang, and Nanni (2001) reported on an in-field fiber reinforced polymer repair of an impact damaged bridge by an over-height vehicle in Missouri. The damaged bridge, Bridge A10062, is located at the interchange of Interstates 44 and 270 in St. Louis County, Missouri. After removal of the loose concrete, inspection of the prestressed concrete bridge girders showed that the exterior girder had sustained the most damage with two of the twenty prestressing strands severed. Figure 2-23 shows some of the damage to the bridge. From an analytical procedure it was determined that a 190 k-ft

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25 moment capacity would be needed from the FRP system to restore the girder to its original strength. The repair called for a two-ply unidirectional carbon fiber reinforcing polymer system with stirrups along the fiber length. No load tests, field measurements, or laboratory simulations were conducted. After the repair was completed, it was inspected and any potential void areas that were identified were epoxy injected to ensure that a complete bond was provided. Scheibel, Parretti, and Nanni (2001) investigated the effectiveness of fiber reinforced polymer strengthening of eleven prestressed concrete girders from another impact damaged bridge in Missouri. The damaged bridge, Bridge A4845, is located over Route 291 on Route 24 in Jackson County, Missouri. An inspection of the eleven girders showed only concrete spalling and the exposure of reinforcing steel and prestressing tendons. Further inspection revealed that only three girders had exposed prestressing tendons with a maximum of two tendons exposed in one girder. Figures 2-24 and 2-25 show overall and detailed views of the damaged bridge. For the design of the strengthening system, the exposed tendons were assumed to be 50% effective. From an analytical procedure it was determined that a 187 k-ft moment capacity would be needed from the FRP system to restore the girder to its original strength. The repair called for a two-ply unidirectional carbon fiber reinforcing polymer system with stirrups along the fiber length. Long-term durability was addressed by applying the repair in an area away from the damage to perform bond pull-off and torsion tests over five years. Initial durability tests indicated a strong bond between the concrete surface and the FRP. No load tests, field measurements, or laboratory simulations were conducted to show that the designed repair was sufficient to restore the girders original load capacity.

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26 Boyd and Banthia (2001) reported on a new application method of fiber reinforced polymer strengthening systems in which the fibers and matrix resin are simultaneously sprayed onto the member surface. The spraying equipment provides a two-dimensional random distribution of fibers with adjustable fiber lengths and has the ability to build up the FRP strengthening system to any desired thickness. To examine the effectiveness of this new method, three reinforced concrete channel beams that had been removed from a badly deteriorated bridge were tested under third point bending to failure. The damage to the beams consisted of cracking, loss of concrete cover, and reinforcement corrosion, which varied, in each specimen making direct comparisons invalid. From the three specimens, one was treated as a control with no repair while the other two were both repaired with E-glass fibers: one repaired with the spray technique and the other repaired with a continuous fiber system. Both applications were completed under laboratory conditions meaning that the beams were repaired while upside down, which is not possible in the field. The spray consisted of E-glass fibers embedded in a matrix consisting of a polyester resin and a catalyst for the resin. The researchers also conducted coupon tests and determined that the strength of the fiber reinforced polymer system increased with increasing fiber length. Figure 2-26 shows the load deflection relationship for the control specimen, the fabric repaired specimen, and the spray repaired specimen. The experiments showed a greater increase in stiffness with the spray repair over the fabric repair and a significant increase in strength for both repair fabric types. Also, for material costs only, the spray repair was determined to be less expensive than the continuous fiber repair.

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27 Spadea, Swamy, and Bencardino (2001) reported on the laboratory testing of reinforced concrete beams repaired with carbon fiber reinforced polymer laminates reporting on respect to ductility, strength and failure modes. Eleven rectangular reinforced concrete beams were constructed. The beams were divided into three series, one control beam for each series with no external reinforcing referred to as A1, A2, and A3. Within a series the internal reinforcing remained constant and the external reinforcing was varied. Beams 1.1 and 3.1 were strengthened longitudinally with one sheet of CFRP and no additional reinforcement. Each series contained a beam (1.2, 2.2, and 3.2) that was strengthened longitudinally with one sheet of CFRP, a wide stirrup at each support, and four small stirrups spaced along the length of the beam. Beams 1.3 and 3.3 were strengthened longitudinally with one sheet of CFRP, a wide stirrup at each support, two small stirrups at midspan, and three small stirrups at each load point with a plate above. Beam 2.3 was strengthened longitudinally with one sheet of CFRP, a wide stirrup at each support, and ten small stirrups spaced along the length of the beam. The stirrups or anchorages were made of steel and applied to the beams with the same adhesive used for the CFRP. The steel plate anchorages were designed to either counteract stresses at the end of the plates and control bond slip between the CFRP plate and concrete, restrain movement of the plate, or provide lateral confinement for the concrete in compression. The beams were tested in four-point bending to failure to determine what increases in strength, ductility, and stiffness could be attained from the different reinforcing schemes. The control beams with no external reinforcing failed as expected due to concrete crushing. Beams 1.1 and 3.1 with only one sheet of CFRP and no anchorages, failed suddenly due to debonding of the CFRP at a capacity 60.7% and

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28 30.8% greater than their corresponding control specimen. The other beams, with longitudinal and varying external stirrup reinforcing, all resisted a higher load than their respective control specimen and failed in a more ductile manner than the beam with only longitudinal reinforcing. The beams with anchorage reinforcing also reached higher strains in the concrete and in the CFRP than the beams with only longitudinal reinforcing. Structural ductility was quantified for deflection, curvature, and energy as defined by shown in Equations 2-1, 2-2, and 2-3. yu (2-1) yu (2-2) ytotEEE (2-3) Where u = midspan deflection at ultimate load, y = midspan deflection at yield, u = curvature at ultimate load, y = curvature at yield load, E tot = area under the curve at ultimate load, and E tot = area under the curve at ultimate load. In Equation 3, E tot is taken as the entire area under the load deflection curve and E y is taken as the area under the load deflection curve up to yield. Table 2-7 provides the defined ductility ratios for all test specimens. Comparing the defined ductility ratios of the control beam to the other beams shows a significant reduction for all of the repaired beams and shows a more significant reduction in ductility of the beams repaired with only longitudinal reinforcing over the beams with longitudinal and stirrup reinforcing.

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29 Table 2-1: Qualitative Comparison of Carbon, Aramid and Glass Fibers (Meier and Winistorfer 1995) Carbon FibersAramid FibersE-Glass FibersTensile StrengthVery GoodVery GoodVery GoodCompressive StrengthVery GoodInadequateGoodYoung's ModulusVery GoodGoodAdequateLong-Term BehaviorVery GoodGoodAdequateFatigue BehaviorExcellentGoodAdequateBulk DensityGoodExcellentAdequateAlkaline ResistanceVery GoodGoodInadequatePriceAdequateAdequateVery GoodFiber Composite Sheets Made of:Criterion Table 2-2: Comparison of Load, Deflection, Stiffness, and Crack Width (Arockiasamy 1995) Beam# of platesUltimate Load (kip)% diffMidspan Deflection at Ultimate (in)% diffInitial Stiffness (kip-in/in)% diffMidspan Crack Width at Ultimate (in)% diffS5-STL013.442-1.65-674.685-0.0147-S5-PRE1114.97311.391.3418.79813.80320.620.0029480S5-PRE2219.87847.880.99439.76980.72545.360.009237.41S6-PRE3222.01463.770.96241.71004.55548.890.0055462.31S6-PRE4320.67553.810.64361.031011.46549.920.0047467.76S6-PRE5326.12694.360.88146.611197.9577.560.0045668.98

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30 Table 2-3: Comparison of Midspan Deflection Before and After FRP Application (Tedesco et al. 1998) GirderBefore FRP (mm)After FRP (mm)Percent Difference165.627.97.3376.343.22.815.45.227.77.137.36.643.73.313.1326.5638.67.746.35.513.63.526.96.338.37.445.65Loading Position 1Loading Position 2Loading Position 3Loading Position 4 781012581012281012491112 Table 2-4: Comparison of Reinforcing Bar Stresses Before and After FRP Application (Tedesco et al. 1998) GirderBefore FRP (MPa)After FRP (MPa)Percent Difference18377291853827443734175722888238477445411393727266310694114827414744276693102901247366Loading Position 1Loading Position 2Loading Position 3Loading Position 4 77109478948106910

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31 Table 2-5: Comparison of Initial Stiffness and Strength for Various Retrofits (Boyd and Banthia 2001) Initial (kN/mm)Change (%)Peak (kN)Change (%)None6.69-214-Fabric7.671528433Spray93541996StiffnessStrengthRetrofit Type Table 2-6: Comparison of Stiffness Before and After Retrofit (Boyd and Banthia 2001) Retrofit TypeStiffness Before Retrofit (kN/mm)Stiffness After Retrofit (kN/mm)Change (%)Fabric6.97.6711Spray7.78916 Table 2-7: Comparison of Ductility Ratios (Spadea et al. 2001) BeamFailure Load (kN)Deflection DuctilityCurvature DuctilityEnergy DuctilityA1546.2811.6A1.186.81.51.51.9A1.2984.32.99.5A1.396.73.547.6A229.27.511.517.4A2.2754.8313.2A2.377.75.8416.1A357.2710.315.7A3.174.822.23.2A3.298.84.53.310.9A3.398.33.84.98.4

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32 Figure 2-1: Stress-Strain Diagrams for Mild Steel and Carbon, Glass, and Aramid Fiber Reinforced Polymers (FIB 2001) Figure 2-2: FRP Placement for Increase in Shear Capacity

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33 Figure 2-3: Flexural Failure Modes for FRP Strengthened Beams (Teng et al. 2002)

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34 (a) (b) Figure 2-4: Shear Failure Modes for FRP Strengthened Beams (Teng et al. 2001) Figure 2-5: Load versus Midspan Deflection for Specimen 1 with 54 ksi yield steel (Sen and Liby 1994)

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35 Figure 2-6: Load versus Midspan Deflection for Specimen 2 with 54 ksi yield steel (Sen and Liby 1994) Figure 2-7: Load versus Midspan Deflection for Specimen 3 with 45 ksi yield steel (Sen and Liby 1994)

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36 Figure 2-8: Load versus Midspan Deflection for Specimen 4 with 54 ksi yield steel (Sen and Liby 1994) Figure 2-9: Load versus Midspan Deflection for Specimen 5 with 45 ksi yield steel (Sen and Liby 1994)

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37 Figure 2-10: Load versus Midspan Deflection for Specimen 6 with 45 ksi yield steel (Sen and Liby 1994)

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38 Figure 2-11: Moment Deflection Relationship for Six Specimens (Arockiasamy 1995)

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39 Figure 2-12: Moment versus CFRP Strain for Repaired Specimens (Arockiasamy 1995)

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40 Figure 2-13: Moment versus Reinforcing Bar Strains for Six Specimens (Arockiasamy 1995)

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41 Figure 2-14: Moment versus Concrete Compressive Strain for Six Specimens (Arockiasamy 1995)

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42 Figure 2-15: Overall Damage to I-680 over CR L34, near Beebeetown, Iowa (Klaiber et al. 1999) Figure 2-16: Moment Deflection Relationships for Beams 1 and 2 (Klaiber et al. 1999)

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43 Figure 2-17: Experimental and Theoretical Load Deflection Relationship for Beams 1 and 2 (Mayo et al. 1999) Figure 2-18: Load Deflection Relationship of Fully-Wrapped Beams (Shahawy and Beitelman 1999)

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44 Figure 2-19: Relationship Between Number of FRP Layers to Moment Capacity at Steel Yield and Ultimate (Shahawy and Beitelman 1999) Figure 2-20: Load Deflection Relationships for Fully-Wrapped and Partially-Wrapped Sections with 3 Layers of FRP (Shahawy and Beitelman 1999)

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45 Figure 2-21: Theoretical and Experimental Midspan Deflection for the Analyzed Sections (Shahawy and Beitelman 1999)

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46 Figure 2-22: Load Deflection Relationship for Beams 1, 2, 3, and 4 (Labossieere et al. 2000) Figure 2-23: Damage to Bridge A10062, St. Louis County, Missouri (Tumialan et al. 2001)

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47 Figure 2-24: Overall Damage to Bridge A4845, Jackson County Missouri (Scheibel et al. 2001) Figure 2-25: Close-up Damage to Bridge A4845, Jackson County Missouri (Scheibel et al. 2001)

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48 Figure 2-26: Load Midspan Deflection Relationship for the Control, Repaired with Fabric, and Repaired with Spray Specimens (Boyd and Banthia 2001)

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CHAPTER 3 DESCRIPTION OF EXPERIMENTAL STUDY In order to test different repair methods several identical test specimens were needed that were in similar conditions. Six unutilized prestressed concrete type II AASHTO bridge girders were found in a pre-cast yard in south Florida. The 43-9 girders were thought to be approximately 26 years old and to have the same constituents based on corresponding lengths, strand patterns, and stirrup locations. These six girders provided the testing of four repair methods leaving two girders as control specimens. Two control specimens were needed to establish the moment and shear capacities at an undamaged and damaged condition and therefore verify that sufficient additional moment and shear strength would be attained from the various repair techniques to restore the original moment and shear capacities of the girder. Before any testing was initiated, a 1-0 deep by 2-0 wide slab having an average 6300 psi compressive strength was cast on each of the girders. The slabs were cast in order to counteract the effect of camber due to prestressing and to duplicate in-field conditions of having a deck or dead load acting on the girders. Figure 3-1 shows a detailed cross-section and profile of the girders in the undamaged state. Description of Test Specimens in Damaged State All of the test specimens except for the undamaged control specimen were damaged to replicate the most probable conditions that impact of an over-height vehicle would cause. To simulate the impact damage in the laboratory, first a 5-0 long section of concrete was removed from each side of the bottom flange centered about midspan. 49

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50 The hatched area in Figure 3-2 represents the removed concrete. After removal of the concrete, four prestressing strands were exposed, two on each side, that were severed representing an 18% loss in capacity. A concrete patch was then applied to the removed concrete area to restore the original gross cross-section. An average 4500 psi concrete compressive strength for all of the girders was determined by testing cores taken from one test specimen. The location of the prestressing hold downs at the third points was determined when the concrete was removed to damage the specimens. Direct tension tests were performed of the removed 7/16 diameter, 250 ksi strands. Figure 3-2 shows the cross-section of the girders at their midspans and ends and gives details as to the damage executed on the girders. Figure 3-3 shows a typical girder in the damaged state before the concrete patch was applied. The results of the direct tensile tests for the removed strands are shown in Figure 3-4. The surfaces of all of the girders were sandblasted on the sides and bottom in order that a good bond surface for the FRP repairs would be produced. Test Setup for Four Point Bending Tests All girders were tested in a four point bending configuration where a uniform moment exists between the load points and shear is only present between the load points and reactions. Figure 3-5 shows the load diagram and resulting shear and moment diagrams for the experiments. For the test setup, the load points were 10-0 apart and the supports were 40-0 apart centered about the midspan. Figure 3-6 shows the test setup for all test specimens and includes the locations of crack gauges, LVDTs, and load cells. Figures 3-6b and 3-6c show the location of instruments corresponding to the instrumentation lines shown in Figure 3-6a. The girders were loaded statically to failure noting crack locations, sizes, and propagations during loading. The data from the tests

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51 was reduced to determine the moment capacity, shear capacity, and midspan deflection of the undamaged and damaged control specimens to be used for comparison and to later verify if the FRP repair of the damaged girders had reached the required capacities. The capacities would also provide the necessary values to be utilized in the design. The supports for Test Specimen 4 were moved out an additional 6 on both ends due to the termination of the FRP laminates occurring at the location of the supports for Tests 1, 2, 3, 5, and 6. Repair Designs and Design Properties The company that was to perform the repair was responsible for the design calculations. No method or procedure, factors of safety, assumption for failure mode, or fiber or resin material type was specified for the repair designs by the FDOT. Appendix A contains the repair designs for Test Specimens 3, 4, and 5. A comparison of the repair design properties and the corresponding description of the repair for Test Specimens 3, 4, 5, and 6 is shown in Table 3-1. The design properties in Table 3-1 for the repairs of Test Specimens 3, 4, and 6 are from data sheets provided by the manufacturers through direct tensile coupon tests. ASTM D3039 (2000) gives standards for determining tensile properties of polymer matrix composite materials. The standards state that the extensometer gage length should be in the range of 0.5 to 2.0 inches. An extensometer gauge length of 0.5 was used for the direct tensile coupon test data for Test Specimen 3. An extensometer gauge length of 2 was used for the direct tensile coupon test data for Test Specimen 4. The design properties in Table 3-1 for the repairs of Test Specimen 5 were taken from previous tensile coupon tests of representative properties in which an extensometer gauge length of approximately 2

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52 inches was utilized (Boyd 2000). For the Test Specimen 6 laminate properties, an average of 0.12 elongation over an 8 long specimen was reported. Repair Application Procedures For Test 3, employees of RJWatson performed the installation procedure in which a wet lay-up procedure was implemented. First a layer of epoxy was applied to the tension face of the girder where the FRP strengthening system would be applied. After the epoxy became tacky, one layer of pre-cut fabric was placed on the tension face of the girder and rolled out to remove voids and prevent wrinkling of the fabric. Epoxy was then applied to the exposed side of the fabric. When the exposed epoxy became tacky the next layer of FRP was applied. This procedure was followed for each successive layer of FRP applied. Figure 3-7 shows the installed repair for Test Specimen 3 in which the fabric direction is oriented parallel to the longitudinal axis of the member. It should be noted that all four layers are terminated at the same location. For Test 4, the installation procedure was performed by Air Logistics in which pre-cut fabric sizes of pre-impregnated resin were applied to the tension face with additional stirrups of fabric placed to resist peeling of the fabric. A primer was first applied to the tension side of the girder. When the primer became tacky, the fabric was unrolled directly onto the surface. After two of the four layers were applied, water was sprayed over the fabric with a garden hose activating the resin. The remaining two layers were then applied and water was again sprayed over the fabric to activate the resin. As each layer was applied, it was necessary to stretch the fabric in order to reduce air voids and wrinkling of the fabric. After all of the layers were applied to the tension face, two layers of bi-directional fabric were applied as stirrups following the above procedure with two bolts placed through the ends into the web. Figure 3-8 shows the installed repair for Test

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53 Specimen 4 in which the fabric direction is oriented parallel to the longitudinal axis of the member and the layers are terminated at varying locations. In addition to longitudinal reinforcement, a two-ply stirrup was installed just within the termination of the second shortest layer to enhance the behavior of the FRP strengthening system. The stirrup fabric was oriented at 0 and 90 as represented in Figure3-8 by the hatched area. Magnum Venus Products (MVP) and graduate students from the University of Florida performed the installation for Test 5. Figure 3-9 shows the equipment used for the spray application procedure. The fibers in a roving form are routed into a chopper mechanism attached to the spray gun. A detailed view of the gun is shown in Figure 3-10. The chopping mechanism has two rollers with adjustable blades to vary fiber lengths as the fiber roving passes through the rollers. The final FRP composite is comprised of a resin, catalyst, and fibers that are fed separately into the spray gun and meet simultaneously beyond the gun either in mid-air or on the application surface shown in Figure 3-11. After an application of spraying was complete to an average thickness of 5 mils, ribbed metal rollers were used to compact the fibers, resin, and catalyst onto the surface, remove air voids, and assist in bonding the material to the surface and underlying layers. The above procedure was followed making each pass in which about 5 mils was sprayed onto the surface and rolled out until the desired thickness was reached. The repair was originally designed with fibers approximately 1.9 inches in length and its representative properties with a laminate thickness of 0.35 inches. While attempting the installation with the 1.9 inch fiber length, it was determined that it would be necessary to reduce the fiber lengths. Since it was known from previous research that reducing fiber length causes a reduction in composite strength, the repair was redesigned (Boyd 2000).

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54 A new fiber length of 1.25 inches was utilized with a desired thickness of 0.50 inches to produce an equivalent strength as specified in the original design. Figure 3-12 shows the installed repair for Test Specimen 5 in which the multidirectional or randomly oriented fibers were applied to the entire bottom of the girder and up 6 on both sides of the bottom flange for 20-0 centered about midspan. In addition a 2-0 wide stirrup of randomly oriented fibers was applied up to the bottom of the top flange at the termination of the repair at each end of the girder to enhance the shear strength of the repair and ensure a good bond of the FRP strengthening system. Difficulties were encountered during the two-day application of the spray FRP. The spray application procedure had not been used in this type of application previously where the spray had to be projected up onto the tension face of the beam simulating in-field conditions with multiple layers having to be applied to reach the desired thickness. In order to apply another layer, the underlying layer must be set such that the additional layer does not pull the underlying layer off due to its self-weight. The FRP thickness measurement device was incapable of piercing through the hardened underlying layers to get accurate measurement of applied thickness so it was not known during or after the application exactly what thickness was actually applied. Therefore, FRP cores were taken from the specimen after loading. Figure 3-13 shows where samples were taken from the cross-section and profile of the specimen. A total of 73 samples were taken from the sides and bottom of the test specimen to evaluate if the desired thickness was reached. The side samples are defined are those taken from the web, and the angled and vertical side of the flange (numbers 1 through 5 and 9 through 13) as shown in Figure 3-13. The 73 samples yielded an average thickness

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55 of 0.479 with a standard deviation of 0.196. Of the 21 samples taken from the tension face of the girder none had a thickness equal to the desired thickness of 0.5. The maximum and minimum thicknesses taken from the tension face of the girder were 0.368 and 0.139, respectively with an average thickness of 0.266 and a standard deviation of 0.065. A total of 52 samples, 26 per side, were taken from the sides of the test specimen with an average thickness of 0.565 with a standard deviation of 0.161. There is a significant difference in thickness between the two sides of the specimen from the two days of spraying in which only one side of the beam was repaired each day. When the sides were evaluated by day, the results of the first day showed an average thickness of 0.506 with a standard deviation of 0.102, and the results of the second day had an average thickness of 0.625 with a standard deviation of 0.188. The average thickness, as previously stated, for all of the 73 specimens was 0.479. This was reasonably close to the desired thickness of 0.50, although it is apparent from the bottom thickness measurements that the desired 0.50 was not reached during the application of the FRP to the tension face of the girder. The application procedure for Test Specimen 6 was performed by Edge Composite employees, in which a wet lay-up procedure was implemented. All surface voids, in which any FRP strengthening system was to be applied, were filled with a West System Filler Epoxy containing micro-fibers to form a flat surface for a maximum bond. Then a layer of resin was applied to the tension face of the beam in which the fabric would be applied. When the resin became tacky the first layer of pre-cut fabric was applied to the girder. A plastic trowel or scraper was used to remove any excess resin and to obtain the desired constant thickness of resin between the layers. Each layer of fabric was rolled out

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56 to remove voids and prevent wrinkling of the fabric. Resin was then applied to the exposed side of the first layer of fabric. The procedure described above was followed for the two successive layers. After the three layers were applied to the tension face, a two ply stirrup was applied at the end of the longitudinal fabric to provide resistance against peeling of the longitudinal fabric. The stirrups were applied to the girder using the above described procedure. Figure 3-13 shows the installed repair for Test Specimen 6 in which the fabric direction is oriented parallel to the longitudinal axis of the member and the layers are terminated at the same location. In addition to the longitudinal reinforcement, a two-ply stirrup was installed just within the termination of the fabric to enhance the behavior of the FRP strengthening system. The stirrup fabric was unidirectional and oriented perpendicular to the length of the girder as represented in Figure3-14 by the striped area.

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57 Table 3-1: Comparison of Design Repairs for Test Specimens 3, 4, 5, and 6 Test Number3456CompanyRJWatsonAir LogisticsUFEdge Com p ositesFiber Material and DescriptionCarbon uni-directional with aramid cross fibersCarbon uni-directional with cross fibersE-Glass multi-directional fibersCarbon uni-directional fibersResin MaterialEpoxyPolyurethanePolyesterEpoxyElongation at Failure1.21%1.05%1.32% 1.43%*1.50%Tensile Modulus (ksi)10500110001711 1522*10100Ultimate Tensile Strength (ksi)12711515.66 15.08*150Total Laminate Thickness (in)0.160.1240.35 0.50*0.069Number of Laminates4413Total Length of Laminates (ft)20 20 20 20 40 32 24 16 2020 20 20 Width of Laminates (in)16121818Stirrup DescriptionN/A2 plies of carbon oriented at 0 and 90, 12 ft from midspan 0.5 inches thick 2 feet wide up to bottom of the flange 2 plies of carbon uni-directional up to bottom flange* As-built properties

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58 Figure 3-1: Typical Profile and Cross-Sections of Test Specimens in Undamaged State

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59 Figure 3-2: Typical Cross-Section Details for Test Specimens 2, 3, 4, 5, and 6 Figure 3-3: Typical Damage for Test Specimens 2, 3, 4, 5, and 6 Showing Removed Concrete and Two Severed Prestressing Strands

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60 Figure 3-4: Stress-Strain Curves for Removed Strands from Test Specimen 2 Figure 3-5: Load, Shear, and Moment Diagrams for All Four-Point Bending Tests

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(a) 61 Figure 3-6: Typical Instrumentation for Four-Point Bending Tests a)Typical Instrumentation Layout for Test Specimens 1, 2, 3, 4, 5, 6

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(b) (c) 62 Figure 3-6(continued): Typical Instrumentation for Four-Point Bending Tests b)Typical Instrumentation Locations for Section A-A Cross-Section in Figure 3-6a; c) Typical Instrumentation Locations for Section B-B Cross-Section in Figure 3-6a

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63 Figure 3-7: Test Specimen 3 Repair Design

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64 Figure 3-8: Test Specimen 4 Repair Design

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65 Figure 3-9: Spray Equipment: Gun (G), Catalyst (C), and Resin (R) Figure 3-10: Detailed View of Spray Gun: Resin (R), Catalyst (C), Spray (S), and Chopper (Ch)

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66 Figure 3-11: Spray Technique Procedure Showing How the Glass Fibers, Resin, and Catalyst for Resin are Combined in Mid-Air

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67 Figure 3-12: Test Specimen 5 Repair Design

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68 Figure 3-13: FRP Sampling Cores for Test Specimen 5

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69 Figure 3-14: Test Specimen 6 Repair Design

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CHAPTER 4 EXPERIMENTAL TEST RESULTS This chapter contains the experimental data for the six, four-point bending tests that were conducted. Chapter 5 will serve as a comparison of the experimental results shown in this chapter to calculated theoretical values. Loading was terminated for each specimen once either load cell showed a reduced reading therefore, no post ultimate data is shown for any of the tests. The total applied load is the sum of the two load cell measurements shown in Figure 3-6a. The moment in the constant moment region is calculated by taking one-half the total applied load and multiplying it by 15-0 for Test Specimens 1, 2, 3, 5, and 6 and 15-6 for Test Specimen 4. Strain and deflection profiles are plotted at 25, 50, 75, 90, and 100 percent of the maximum load attained during each experiment. Strain gauges that either were initially defective, or after a certain load level no longer functioned properly, were omitted from the strain profiles. Large increases in the strain readings for a particular gauge can be attributed to crack formation occurring inside of the two-inch gauge length causing increased strain due to stress redistribution. The experimental neutral axis depths were determined by assuming a best-fit linear strain distribution through the cross-section for a specified load level using all properly functioning strain gauges. In order to compare strain values at a certain load level and strain increases between load levels, the strain gauge closest to the prestressing steel, 45 inches down from the top of the slab, was chosen for comparison between all test specimens. In a case where the strain value was unavailable at the steel level the value was determined by linear interpolation between the 70

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71 compressive strain measured 6 down from the top of the slab and the extreme tensile strain measured 48 down from the top of the slab. If the extreme tensile strain value at 48 down from the top of the slab was also unavailable, the strain at 48 down was set equal to the bottom strain being measured closest to the unavailable data at the same load level. Test 1 Undamaged Control Results The load versus midspan deflection for Test Specimen 1 is shown in Figure 4-1. The experimental load capacities at cracking and ultimate are 131 kips and 268 kips with corresponding midspan deflections of 0.50 and 5.95 inches, respectively. The load deflection relationship, as shown in Figure 4-1, is linear up to the cracking load and nonlinear between the cracking load and maximum load. Symmetrical deflections occurred about the midspan of the beam as shown in Figure 4-2, except for a slightly larger deflection at the south load point at the maximum load, P max Figure 4-2 also shows similar increases in deflection between 0 and 25 percent and between 25 and 50 percent of the maximum load, with larger increases in deflection between 50 and 75 percent and between 75 and 90 percent of the maximum load, and a substantial increase in deflection between 90 and 100 percent of the maximum load. The strain profiles in Figures 4-3, 4-4, and 4-5 all show an upward movement of the neutral axis from the initial location of 20.7 inches down from the top of the slab into the slab at P max The final depths of the neutral axis at the north load point, midspan, and south load point locations are 11.27, 9.05, and 5.73, respectively, down from the top of the slab at the maximum load. The variation in neutral axis depths can be attributed to either a slight imbalance of 4 kips at the maximum applied load with a higher load at the south load point location, larger deflections, slipping of prestressing strands, yielding of

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72 the strands, or a compression failure all on the south end of the girder. Figures 4-3, 4-4, and 4-5 also show a significant increase in strain between 90 and 100 percent of the maximum load that can be attributed to the formation of multiple cracks or the further opening up of existing cracks that developed under the applied load. Figure 4-3 shows only slight increases in strain when compared to the strain increases of Figures 4-4 and 4-5. Figures 4-4 and 4-5 show similar strain profiles for the midspan and south load point, which is reasonable since both locations are in the constant moment region. Strain increases between the load levels at the prestressing steel layer at the north load point shown in Figure 4-3 are 245 x 10 -6 in/in, 1741 x 10 -6 in/in, 1933 x 10 -6 in/in, and 3508 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at midspan shown in Figure 4-4 are 456 x 10 -6 in/in, 3612 x 10 -6 in/in, 2615 x 10 -6 in/in, and 7240 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at the south load point shown in Figure 4-5 are 339 x 10 -6 in/in, 3303 x 10 -6 in/in, 2224 x 10 -6 in/in, and 7680 x 10 -6 in/in. Comparing the strain increases at the prestressing steel level shows significantly larger increases in strain at the south load point than at the north load point, and strain increases at the south load point similar to the increases at midspan. The larger strains at the south load point location can be attributed to either unbalanced load cells with a higher load at the south load point location, larger deflections, slipping of prestressing strands, yielding of the strands, or a compression failure all on the south end of the girder. Figure 4-6 shows the load versus strain data for the strain gauges shown in Figure 3-6a along the bottom centerline of the test specimen. The figure shows similar strain readings for the four gauges up to the cracking load, and similar strains up to failure for the two strain gauges at 1-0 from midspan and for the two strain gauges

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73 3-0 from midspan. The maximum recorded bottom centerline strain at P max was 0.018447 in/in and occurred at 3-0 north of midspan. Concrete crushing was not the failure mode as expected, which allows for slipping of the strands to be assessed as the failure mode. The slightly larger deflections and larger strains recorded at the south load point location provide an assumption that the slipping of strands occurred on the south end of the girder. The shear, moment, and midspan deflection recorded at failure for the undamaged girder were 268 kips, 2012 kip-ft, and 5.95 inches, respectively. Test 2 Damaged Control Results The load versus midspan deflection for Test Specimen 2 is shown in Figure 4-7. The experimental load capacities at cracking and ultimate are 112 kips and 210 kips with corresponding midspan deflections of 0.54 and 6.42 inches, respectively. The load deflection relationship, as shown in Figure 4-7, is linear up to the cracking load and nonlinear between the cracking load and maximum load. Symmetrical deflections occurred about the midspan of the beam as shown in Figure 4-8, except for a slightly larger deflection at the south load point at the maximum load, P max Figure 4-8 also shows similar increases in deflection between 0 and 25 percent and between 25 and 50 percent of the maximum load, with larger increases in deflection between 50 and 75 percent and between 75 and 90 percent of the maximum load, and a substantial increase in deflection between 90 and 100 percent of the maximum load. The strain profiles in Figures 4-9, 4-10, and 4-11 all show an upward movement of the neutral axis from the initial location of 20.7 inches down from the top of the slab into the slab at the maximum load. The final depths of the neutral axis at the north load point, midspan, and south load point locations are 10.78, 8.16, and 5.33, respectively, down

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74 from the top of the slab at the maximum load. The variation in neutral axis depths can be attributed to either a slight imbalance of 1 kip at the maximum applied load with a higher load at the south load point location, larger deflection, slipping of prestressing strands, yielding of the strands, or a compression failure all on the south end of the girder. Figures 4-9, 4-10, and 4-11 also show a significant increase in strain between 90 and 100 percent of the ultimate load possibly due to the formation of multiple cracks or the further opening up of existing cracks. The strain profiles in Figures 4-9 and 4-11 for the north and south load point locations show similar strain values for all load increments. A comparison of the strains in Figures 4-9, 4-10, and 4-11 show that there were larger increases at midspan than at the load point locations for all profiles up to 0.90P max This can be attributed to the damage imparted to the specimen only in the midspan area prior to testing. Strain increases between the load levels at the prestressing steel layer at the north load point shown in Figure 4-9 are 254 x 10 -6 in/in, 1364 x 10 -6 in/in, 1121 x 10 -6 in/in, and 11759 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at midspan shown in Figure 4-10 are 536 x 10 -6 in/in, 6392 x 10 -6 in/in, 16971 x 10 -6 in/in, and 1140 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at the south load point shown in Figure 4-11 are 277 x 10 -6 in/in, 1243 x 10 -6 in/in, 800 x 10 -6 in/in, and 11832 x 10 -6 in/in. Comparing the strain increases at the prestressing steel level shows larger increases in strain at the midspan location than at either of the load point locations except for the last load step where similar increases in strain at the north and south load point locations were recorded. Larger increases at the midspan location than at the load point locations are reasonable since the midspan is within the constant or maximum moment region and the load points are not located in the

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75 constant moment region. Figure 4-12 shows the load versus strain for the strain gauges shown in Figure 3-6a along the bottom centerline of the test specimen. Similar strain readings were recorded up to the cracking load for the four strain gauges. The two strain gauges at 1-0 from the midspan show similar strains up to failure and the two strain gauges 3-0 from midspan show similar strains up to failure. Figure 4-12 also shows higher strains for the strain gauges 1-0 from midspan at the maximum load than the strain gauges 3-0 from midspan, which were located one foot from the end of the damaged area. The maximum recorded bottom centerline strain at P max was 0.023338 in/in and occurred at 1-0 south of midspan. Concrete crushing was not the failure mode as expected, which allows for slipping of the strands to be assessed as the failure mode. The slightly larger deflections and larger strains recorded on the south load point location provide an assumption that the slipping of strands occurred on the south end of the girder. The shear, moment, and midspan deflection recorded at failure for the damaged girder were 210 kips, 1576 kip-ft, and 6.42 inches, respectively. Test 3 RJWatson Repair Results Two cycles of loading and unloading of Test Specimen 3 were performed before the specimen was taken to failure as shown in Figure 4-13. The data provided in Figures 4-15 through 4-18 was taken from the last series of loading. The two initial cycles of loading caused a slight softening if the load deflection curve, meaning that there was some measurable reduction in the stiffness of the specimen at the beginning of the loading cycle to failure. The load versus midspan deflection for Test Specimen 3 is shown in Figure 4-14. The experimental load capacities at cracking and ultimate are 125 kips and 244 kips with corresponding midspan deflections of 0.60 and 3.15 inches,

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76 respectively. The experimental cracking load for Test Specimen 3 was attained during the second series of loading which means that for the third or final loading series, the cracking load had previously been exceeded. The load deflection relationship, as shown in Figure 4-14 is still linear up to the cracking load and nonlinear between the cracking load and maximum load. Symmetrical deflections occurred about the midspan of the beam as shown in Figure 4-15, except for a slightly larger deflection at the south load point at the maximum load. Figure 4-15 also shows similar increments of deflection between 0 and 25 percent and between 25 and 50 percent of the maximum load, with increasingly larger increments of deflection between 50 and 75 percent 75 and 90 percent, and 90 and 100 percent of the maximum load. The strain profiles in Figures 4-16, 4-17, and 4-18 all show an upward movement of the neutral axis from the initial location of 20.81 inches down from the top of the slab into the slab at P max The final depths of the neutral axis at the north load point and midspan locations are 10.98 and 10.52, respectively, down from the top of the slab at the maximum load. The neutral axis depth at the south load point location can not be determined due to an insufficient number of data points. The determined final depths of the neutral axis for the north load point and midspan locations are similar with a slightly smaller neutral axis depth for the midspan location. This difference can be attributed to either the midspan location being within the constant or maximum moment region and the north load point location being located at the transition between the moment gradient and constant moment regions or a slight imbalance of 7 kips at the maximum applied load with a higher load at the south load point location. The strain profiles in Figures 4-9 and 4-11 for the north and south load point locations show similar strain patterns for all

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77 load steps with slightly larger strains at the north load point. A comparison of the strains in Figures 4-9, 4-10, and 4-11 show that there were larger increases at midspan than at the load point locations for all profiles. This can be attributed to the damage imparted on the specimen only in the midspan area prior to testing. Strain increases between the load levels at the prestressing steel layer at the north load point shown in Figure 4-16 are 254 x 10 -6 in/in, 1610 x 10 -6 in/in, 1210 x 10 -6 in/in, and 753 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at midspan shown in Figure 4-17 are 503 x 10 -6 in/in, 1454 x 10 -6 in/in, 1118 x 10 -6 in/in, and 1174 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at the south load point shown in Figure 4-18 are 209 x 10 -6 in/in, 810 x 10 -6 in/in, 650 x 10 -6 in/in, and 2347 x 10 -6 in/in. The largest increases for the strain gauges near the level of prestressing were located at midspan for the first two load steps of 0.25 and 0.50P max the north load point for the next two load steps of 0.75 and 0.90P max and the south load point for the last load step up to P max The pattern of strain increase shows fairly symmetric strains about midspan up to 0.50P max equal to 122 kips which is approximately equal to the experimental cracking load of 125.28 kips. Also, the largest strain increase occurred between 0.90P max and 1.00P max at the south load point, which is at the location where a portion of the girder completely separated from the rest of girder at failure. Figure 4-19 shows the load versus strain data for the strain gauges shown in Figure 3-6a along the bottom centerline of the test specimen. Similar strain readings were recorded for the two strain gauges on the north end of the beam up to failure. The figure shows slightly larger strains for the strain gauges located 3-0 from midspan than the strain gauges 1-0 from

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78 midspan throughout loading. The maximum recorded bottom centerline strain at P max was 0.005083 in/in and occurred at 3-0 north of midspan. The failure mode of Test Specimen 3 was concrete cover separation as shown in Figures 4-20 and 4-21. A crack formed at both ends of the FRP plate at the bottom layer of prestressing and propagated towards midspan. The severe cracking at the south end of the girder caused a portion of the bottom flange of the girder up to the bottom layer of prestressing to fall to the ground with the adhered FRP. The shear, moment, and midspan deflection recorded at failure for the repaired girder were 244 kips, 1829 kip-ft, and 3.15 inches, respectively. The ultimate shear and moment capacity corresponds to restoring 91% of the experimental ultimate capacity of the undamaged girder. The measured deflection at P max for Test Specimen 3 corresponds to a 47% reduction in the deflection of the undamaged girder (Test Specimen 1) at its maximum load. Test 4 Air Logistics Repair Results The load versus midspan deflection for Test Specimen 4 is shown in Figure 4-22. The experimental load capacities at cracking and ultimate are 111 kips and 202 kips with corresponding midspan deflections of 0.47 and 1.90 inches, respectively. The load deflection relationship, as shown in Figure 4-22 is linear up to the cracking load and nonlinear between the cracking load and maximum load. Symmetrical deflections occurred about the midspan of the beam as shown in Figure 4-23, except for a slightly larger deflection at the south load point at the maximum load. Figure 4-23 also shows similar increases in deflection between 0 and 25 percent and between 25 and 50 percent of the maximum load, with slightly larger increases in deflection between 50 and 75 percent 75 and 90 percent, and 90 and 100 percent of the maximum load.

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79 The strain profiles in Figures 4-24, 4-25, and 4-26 all show an upward movement of the neutral axis from the initial location of 20.76 inches down from the top of the slab into the slab at P max The final depths of the neutral axis at the north load point, midspan, and south load point locations are 9.08, 8.18, and 14.62, respectively, down from the top of the slab at the maximum load. It is reasonable that the neutral axis depth at midspan would be less than the neutral axis depths at the load point locations due to the midspan location being within the constant or maximum moment region and the load point locations being located at the transition between the moment gradient and constant moment regions. The difference of the neutral axis depths at the load point locations can be attributed to either a slight imbalance of 1 kip at the maximum applied load with a higher load at the north load point location, larger deflections, slipping of prestressing strands, yielding of the strands, compression failure, or failure of the FRP all on the north end of the girder. The strain profiles in Figures 4-24 and 4-26 for the north and south load points show similar strain patterns for all load increments with slightly larger strains at the north load point. A comparison of the strains in Figures 4-24, 4-25, and 4-26 show that there were larger increases at midspan than at the load points for all profiles. This can be attributed to the damage imparted to the specimen only in the midspan area prior to testing. Strain increases between the load levels at the prestressing steel layer at the north load point shown in Figure 4-24 are 157 x 10 -6 in/in, 862 x 10 -6 in/in, 1146 x 10 -6 in/in, and 794 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at midspan shown in Figure 4-25 are 802 x 10 -6 in/in, 1121 x 10 -6 in/in, 1158 x 10 -6 in/in, and 1103 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at the south load point shown in Figure 4-26 are 122 x 10 -6 in/in,

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80 1572 x 10 -6 in/in, 959 x 10 -6 in/in, and 931 x 10 -6 in/in. The largest increases for the strain gauges near the level of prestressing were located at midspan for the first load step up to 0.25P max the south load point for the next load step up to 0.50P max and the midspan for the last two load steps up to P max The strain increase pattern shows overall symmetric strain increases about midspan except between 0.50 and 0.75P max Figure 4-27 shows the load versus strain data for the strain gauges shown in Figure 3-6a along the bottom centerline of the test specimen. Similar strain readings were recorded for the four strain gauges up to the maximum load. Figure 4-27 also shows slightly larger strains at the maximum load for the strain gauges located on the north end of the girder than for the strain gauges located on the south end of the girder. The maximum recorded bottom centerline strain at P max was 0.004897 in/in and occurred at 1-0 north of midspan. The failure mode of Test Specimen 4 was a combination of FRP rupture and adhesive failure as shown in Figure 4-28. Before failure, the FRP system had debonded between the stirrups and was stretched in tension due to the deflection of the girder. At failure, the FRP had reached its tensile capacity in which the FRP ruptured just before the stirrup at the south end and debonded past the stirrup at the north end. By debonding past the stirrup, the longitudinal FRP sheared through the stirrup, which had fibers oriented at 0 and 90 as shown in Figure 4-29. The shear, moment, and midspan deflection recorded at failure for the repaired girder were 202 kips, 1562 kip-ft, and 1.902 inches, respectively. The ultimate shear and moment capacity corresponds to restoring 75% and 78%, respectively, of the experimental ultimate capacity of the undamaged girder. The restored percentages for shear and moment capacity differ due to the supports being moved out an additional 6 for the test as shown in Figure 3-5. The measured deflection

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81 at P max for Test Specimen 4 corresponds to a 68% reduction in the deflection of the undamaged girder at its maximum load. Loading was discontinued for Test Specimen 4 when the FRP system failed at a load corresponding to less than the ultimate experimental load of Test Specimen 2, representing the girder in the damaged state. It can be assumed that Test Specimen 4 would have ultimately failed at approximately the ultimate load of Test Specimen 2. Test 5 UF Sprayed Repair Results The load versus midspan deflection for Test Specimen 5 is shown in Figure 4-30. The experimental load capacities at cracking and ultimate are 125 kips and 255 kips with corresponding midspan deflections of 0.53 and 3.16 inches, respectively. The load deflection relationship, as shown in Figure 4-30 is linear up to the cracking load and nonlinear between the cracking load and maximum load. Symmetrical deflections occurred about the midspan of the beam as shown in Figure 4-31, except for a slightly larger deflection at the north load point at the maximum load. Figure 4-31 also shows similar increases in deflection between 0 and 25 percent and between 25 and 50 percent of the maximum load, with slightly larger increases in deflection between 50 and 75 percent 75 and 90 percent, and 90 and 100 percent of the maximum load. The strain profiles in Figures 4-32, 4-33, and 4-34 all show an upward movement of the neutral axis from the initial location of 20.96 inches down from the top of the slab into the slab at P max The final depths of the neutral axis at the north load point, midspan, and south load point locations are 16.05, 9.18, and 9.12, respectively, down from the top of the slab at the maximum load. The significant difference between the neutral axis depths at the load point locations can be attributed to either a slight imbalance of 6 kips at the maximum applied load with a higher load at the south load point location, larger

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82 deflections, slipping of prestressing strands, yielding of the strands, compression failure, or failure of the FRP all on the south end of the girder. The strain profiles in Figures 4-32 and 4-34 for the north and south load point locations show similar strain patterns for all load increments with slightly larger strains at the south load point. A comparison of the strains in Figures 4-32, 4-33, and 4-34 show that there were larger increases at midspan than at the load point locations for all profiles. This can be attributed to the damage imparted to the specimen only in the midspan area prior to testing. Strain increases between the load levels at the prestressing steel layer at the north load point shown in Figure 4-32 are 254 x 10 -6 in/in, 2084 x 10 -6 in/in, 1451 x 10 -6 in/in, and 952 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at midspan shown in Figure 4-33 are 808 x 10 -6 in/in, 2714 x 10 -6 in/in, 2487 x 10 -6 in/in, and 2814 x 10 -6 in/in. Strain increases between the load levels at the prestressing steel layer at the south load point shown in Figure 4-34 are 209 x 10 -6 in/in, 1149 x 10 -6 in/in, 1733 x 10 -6 in/in, and 1303 x 10 -6 in/in. Comparing the strain increases at the prestressing steel level shows larger increases in strain at the midspan location than at either of the load point locations and similar increases in strain at the north and south load point locations. These larger increases at the midspan location are reasonable since the midspan is within the constant or maximum moment region and the load points are located outside the constant moment region. Figure 4-35 shows the load versus strain data for the strain gauges shown in Figure 3-6a along the bottom centerline of the test specimen. Similar strain readings were recorded for the two strain gauges at 1-0 from midspan up to failure and for the two strain gauges 3-0 from midspan up to failure. In Figure 4-35, the strain gauge located 3-0 from midspan on the south end of the girder is experiencing increased

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83 sensitivity noted by the repeated back and forth motion although the gauge remained functional throughout the loading. Figure 4-35 also shows larger strains after the cracking load for the strain gauges 1-0 from midspan than the strain gauges 3-0 from midspan which are located one foot from the end of the damaged area. The maximum recorded bottom centerline strain at P max was 0.009234 in/in and occurred at 1-0 north of midspan. The failure mode of Test Specimen 5 was FRP rupture as shown in Figures 4-36 and 4-37. A crack formed near midspan of the beam rupturing completely through the FRP applied to the perimeter of the section. The FRP remained attached over the entire length and perimeter of the girder. The shear, moment, and midspan deflection recorded at failure for the repaired girder were 255 kips, 1911 kip-ft, and 3.16 inches, respectively. The ultimate shear and moment capacity corresponds to restoring 95% of the experimental ultimate capacity of the undamaged girder. The deflection at P max for Test Specimen 5 corresponds to a 47% reduction in the deflection of the undamaged girder at its maximum load. Test 6 Edge Composites Repair Results Test Specimen 6 was not tested at this time. Evaluation of Experimental Observations Table 4-1 shows the capacities, midspan deflections, strains, and initial flexural stiffness at the cracking load and Table 4-2 shows the capacities, midspan deflections, and strains at the maximum load. Effect of Fiber Reinforced Polymers on Stiffness A significant reduction in stiffness occurred when four prestressing strands were severed in the test specimens to simulate the impact damage. This stiffness loss resulted

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84 in larger deflections at reduced load levels. The maximum load and subsequent failure also occurred at reduced load levels. The experimental initial flexural stiffness of the test specimens was determined by dividing the load at cracking by the deflection at cracking. The initial flexural stiffness calculated for Test Specimen 1 was 262 kips/in while Test Specimen 2 had an initial flexural stiffness of 208 kips/in as shown in Table 4-1. Therefore, a 21% reduction in initial flexural stiffness was simulated by the severing of four prestressing strands as shown in Table 4-1. The application of FRP materials to the tension face of a girder provides resistance to deflection through the high tensile strength of the FRP materials in comparison to concrete or mild steel. The resulting amount of resistance or stiffness increase is dependent on FRP laminate properties, FRP location on the girder, thickness of the FRP laminate, and an effective bond. Test Specimen 3, repaired with carbon fibers and an epoxy adhesive had an initial stiffness that was only of 79.5% of Test Specimen 1; Test Specimen 4, repaired with carbon fibers with a polyurethane adhesive had an initial stiffness that was 89.4% of Test Specimen 1; and Test Specimen 5, repaired with glass fibers with a polyester adhesive had a comparable initial stiffness that was 89.5% of Test Specimen 1 as shown in Table 4-1. The resulting initial flexural stiffness increases of Test Specimens 3, 4, and 5 over Test Specimen 2 were 0%, 12.8%, and 12.9%, respectively, as shown in Table 4-1. At the cracking load the experimentally measured midspan deflections of Test Specimens 2, 3, 4, and 5 varied in comparison to the measured midspan deflection of Test Specimen 1. The measured midspan deflection at cracking load for Test Specimen 1 was 0.50 while Test Specimen 2 had a midspan deflection at cracking load of 0.54 as shown

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85 in Table 4-1. Therefore, a 7.6% increase in midspan deflection at cracking load was simulated by the severing of four prestressing strands. Test Specimen 3, with a midspan deflection at cracking load equal to 0.60, had a midspan deflection at cracking load 20% greater than Test Specimen 1; Test Specimen 4, with a midspan deflection at cracking load equal to 0.47, had a midspan deflection at cracking load 5.3% less than Test Specimen 1; and Test Specimen 5, with a midspan deflection at cracking load of 0.53, had a midspan deflection at cracking load 6.6% greater than Test Specimen 1 as shown in Table 4-1. The measured midspan deflection at maximum load for Test Specimen 1 was 5.95 while Test Specimen 2 had a midspan deflection at cracking load of 6.42 as shown in Table 4-2. Therefore, a 8% increase in midspan deflection at maximum load was simulated by the severing of four prestressing strands. Test Specimen 3, with a midspan deflection at maximum load equal to 3.15, had a midspan deflection at maximum load 47% less than Test Specimen 1; Test Specimen 4, with a midspan deflection at maximum load equal to 1.90, had a midspan deflection at maximum load 68% less than Test Specimen 1; and Test Specimen 5, with a midspan deflection at maximum load of 3.16, had a midspan deflection at maximum load 47% less than Test Specimen 1 as shown in Table 4-2. Effect of Fiber Reinforced Polymers on Capacity A significant reduction in capacity occurred when four prestressing strands were severed in the test specimens to simulate impact damage. This loss in capacity resulted in failures at reduced load levels. The shear and moment capacity of Test Specimen 1 at the cracking load was 131 kips and 984 kip-ft while the shear and moment capacity of Test Specimen 2 at the cracking load was 112 kips and 839 kip-ft as shown in Table 4-1. Therefore, a 15% reduction in the shear and moment capacity at the cracking load was

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86 simulated by the severing of four prestressing strands. The experimental capacity at the cracking load for Test Specimen 3 was 125 kips which was equal to 96% of the capacity of Test Specimen 1, and equal to 112% gain over the capacity of Test Specimen 2 as shown in Table 4-1. The experimental capacity at the cracking load for Test Specimen 4 was 111 kips which was equal to 85% of the capacity of Test Specimen 1, and equal to 99% of the capacity of Test Specimen 2 as shown in Table 4-1. The experimental capacity at the cracking load for Test Specimen 5 was 125 kips which was equal to 95% of the capacity of Test Specimen 1, and equal to 112% gain over the capacity of Test Specimen 2 as shown in Table 4-1. The shear and moment capacity of Test Specimen 1 at the maximum load was 268 kips and 2012 kip-ft, respectively. While the shear and moment capacity of Test Specimen 2 at the maximum load was 210 kips and 1576 kip-ft, respectively as shown in Table 4-2. Therefore, a 22% reduction in the shear and moment capacity at the maximum load was simulated by the severing of four prestressing strands. The experimental capacity at the maximum load for Test Specimen 3 was 244 kips which was equal to 91% of the capacity of Test Specimen 1, and equal to 116% over the capacity of Test Specimen 2 as shown in Table 4-2. The experimental capacity at the maximum load for Test Specimen 4 was 202 kips which was equal to 75% of the capacity of Test Specimen 1, and equal to only 96% of the capacity of Test Specimen 2 as shown in Table 4-2. The experimental capacity at the maximum load for Test Specimen 5 was 255 kips which was equal to 95% of the capacity of Test Specimen 1, and equal to a 121% over the capacity of Test Specimen 2 as shown in Table 4-2. The amount of capacity increase that the FRP can provide to a structural member is dependent on FRP

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87 laminate properties, FRP location on the girder, thickness of the FRP laminate, and an effective bond. Effect of Fiber Reinforced Polymers on Strains on Bottom Centerline of Test Specimens As a result of smaller midspan deflections of the girders repaired with FRP systems than the unrepaired test specimens, strains at the extreme tensile face of the girder at maximum load were also reduced. The measured tensile strain for Test Specimen 1 at maximum load was 0.018447 in/in while the measured tensile strain for Test Specimen 2 at maximum load was 0.023338 in/in as shown in Table 4-2. Therefore, a 26.5% reduction in measured tensile strain was simulated by the severing of four prestressing strands. Test Specimen 3, with a tensile strain equal to 0.005083 in/in at maximum load, had a tensile strain at maximum load that was 72.5% of Test Specimen 1; Test Specimen 4, with a tensile strain equal to 0.004897 in/in at maximum load, had a tensile strain at maximum load that was 73.5% of Test Specimen 1; and Test Specimen 5, with a tensile strain equal to 0.009234 in/in, had a tensile strain at maximum load that was only 49.9% of Test Specimen 1 as shown in Table 4-2. Other Effects Initial Imperfections and Bond A rough transition from the patch material to the original girder was present on all of the test specimens especially on the sides of the girder. This rough transition made it difficult to obtain a even bond of the FRP material in these areas and could cause high stress concentrations in surrounding areas. This imperfection caused problems for the repair application on Test Specimen 5 due to a significant amount of the repair material being placed on the sides of the girder. The rough transition did not affect the resulting

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88 capacities or failure modes of the test specimens since no significant cracking was present in the FRP near these areas. Before testing the repaired test specimens, an infrared camera was used to scan over the repaired areas of Test Specimens 3, 4, and 5 to determine if existing voids in the bond resulted from the application procedure or from curing. In order to locate voids in the repaired areas, a heat source was passed over the repaired areas. Then when infrared images are taken, a void location is determined by a hotspot in which the heat is collecting. The infrared scanning of Test Specimen 3 showed that no voids were present in the bond at the surface or between the applied layers. The infrared camera scanning did not result in any usable results for Test Specimen 4 since the polyurethane adhesive behaved as an insulator to the heat. Therefore, the FRP material bonded in one area and unbonded in a separate area showed that both locations were unbonded. The infrared scanning of Test Specimen 5 showed the presence of numerous small voids at the exterior surface of the FRP located on the tension face of the girder. Summary of Test Specimen Failure Modes Steel yielding followed by concrete crushing is the expected failure mode for any flexural member. Concrete crushing results when the capacity of the steel is exceeded and therefore the neutral axis depth is reduced to the top of the member such that the section can withstand the applied load in tension. Test Specimens 1 and 2 failed due to slipping of the prestressing strands, not as concrete crushing as expected. Test Specimen 3 failed due to concrete cover separation. Concrete cover seperation is a premature failure, which means that if the ends of the FRP had been properly anchored to the girder a higher capacity would have been attained (Teng et al 2002). At failure of Test Specimen 3, a portion of the bottom flange was completely separated from

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89 the girder. This piece of concrete that was still adhered to the FRP resulted from the concrete being only attached to the girder through the FRP. The failure mode of Test Specimen 4 was FRP rupture following an adhesive failure, which is a premature failure mode meaning that further capacity could have been gained if the FRP material had been applied with a better adhesive. In addition, at failure of Test Specimen 4, the FRP material was no longer adhered to the concrete and was hanging from the bottom of the girder at one end. The resulting failure states of Test Specimens 3 and 4 are not desirable. Overhanging materials with the possibility of falling onto a roadway system, like the failure modes of Test Specimens 3 and 4 would be very dangerous. Test Specimen 5 failed due to FRP rupture, which is not a premature failure mode. Therefore, the full capacity of the installed FRP system was reached at the failure load.

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Table 4-1: Experimental Capacities, Deflections, Strains, and Initial Flexural Stiffness of All Test Specimens at Cracking Load Test SpecimenPcr (kip)% of Test Specimen 1% gain of Test Specimen 2Mcr (kip-ft)cr (in)% loss of Test Specimen 1cr (in/in) maximum in profileLocationof maximumcr in profilecr (in/in) maximum on bottomLocationof maximumcr on bottomInitial stiffness (kip/in)% of Test Specimen 11131.30.0N/A984.40.500.00.000626S 48" down0.000664N 3'262.0100.02111.914.70.0839.40.54-7.60.001838C 48" down0.000597S 1'207.679.23125.34.611.9939.60.60-20.00.000569S 48" down0.000901N 3'208.379.54111.015.4-0.8860.50.475.30.00122C 45" down0.000691N 1'234.189.45125.24.611.8938.70.53-6.60.00103C 48" down0.000824S 1'234.489.56NANANANANANANANANANANANA Table 4-2: Experimental Capacities, Deflections, and Strains of All Test Specimens at Maximum Load Loc Test SpecimenPmax (kip)% of Test Specimen 1% gain of Test Specimen 2Mmax (kip-ft)max (in)% loss of Test Specimen 1max(in/in) maximum in profileLocationof maximum max in profilemax(in/in) maximum on bottomationof maximum max on bottom1268.30.0N/A2012.15.950.00.021053S 48" down0.018447N 3'2210.121.70.01575.96.42-8.00.025185N 48" down0.023338S 1'3243.99.116.11829.13.1547.10.004207S 45" down0.005083N 3'4201.624.9-4.11562.41.9068.00.00522C 45" down0.004897N 1'5254.85.021.31910.83.1646.90.00911C 45" down0.009234N 1'6NANANANANANANANANANA 90

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91 Figure 4-1: Test Specimen 1 Load versus Midspan Deflection Curve Figure 4-2: Test Specimen 1 Measured Deflection Profile

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92 Figure 4-3: Test Specimen 1 Strain Profile at North Load Point Figure 4-4: Test Specimen 1 Strain Profile at Midspan

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93 Figure 4-5: Test Specimen 1 Strain Profile at South Load Point Figure 4-6: Test Specimen 1 Load versus Measured Strain Along Bottom Centerline of Beam

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94 Figure 4-7: Test Specimen 2 Load versus Midspan Deflection Curve Figure 4-8: Test Specimen 2 Measured Deflection Profile

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95 Figure 4-9: Test Specimen 2 Strain Profile at North Load Point Figure 4-10: Test Specimen 2 Strain Profile at Midspan

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96 Figure 4-11: Test Specimen 2 Strain Profile at South Load Point Figure 4-12: Test Specimen 2 Load versus Measured Strain Along Bottom Centerline of Beam

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97 Figure 4-13: Test Specimen 3 Load versus Time Showing Loading and Unloading Cycles Figure 4-14: Test Specimen 3 Load versus Midspan Deflection Curve

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98 Figure 4-15: Test Specimen 3 Measured Deflection Profile Figure 4-16: Test Specimen 3 Strain Profile at North Load Point

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99 Figure 4-17: Test Specimen 3 Strain Profile at Midspan Figure 4-18: Test Specimen 3 Strain Profile at South Load Point

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100 Figure 4-19: Test Specimen 3 Load versus Measured Strain Along Bottom Centerline of Beam Figure 4-20: Test Specimen 3 During Loading Showing Crack Propagation on South End of Girder at the Bottom Layer of Prestressing

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101 Figure 4-21: Test Specimen 3 at Failure Showing Concrete Cover Separation at South End Figure 4-22: Test Specimen 4 Load versus Midspan Deflection Curve

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102 Figure 4-23: Test Specimen 4 Measured Deflection Profile Figure 4-24: Test Specimen 4 Strain Profile at North Load Point

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103 Figure 4-25: Test Specimen 4 Strain Profile at Midspan Figure 4-26: Test Specimen 4 Strain Profile at South Load Point

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104 Figure 4-27: Test Specimen 4 Load versus Measured Strain Along Bottom Centerline of Beam Figure 4-28: Test Specimen 4 at Failure Showing Separation of FRP

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105 Figure 4-29: Test Specimen 4 at Failure Showing Shearing of FRP Stirrup Figure 4-30: Test Specimen 5 Load versus Midspan Deflection Curve

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106 Figure 4-31: Test Specimen 5 Measured Deflection Profile Figure 4-32: Test Specimen 5 Strain Profile at North Load Point

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107 Figure 4-33: Test Specimen 5 Strain Profile at Midspan Figure 4-34: Test Specimen 5 Strain Profile at South Load Point

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108 Figure 4-35: Test Specimen 5 Load versus Measured Strain Along Bottom Centerline of Beam Figure 4-36: Test Specimen 5 at Failure Showing Rupture of FRP at Midspan

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109 Figure 4-37: Test Specimen 5 at Failure Showing Rupture of FRP at Midspan on Tensile Face of Girder

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CHAPTER 5 ANALYTICAL TEST RESULTS This chapter provides a description of the methods used to calculate theoretical capacities of the test specimens and a comparison between the calculated theoretical capacities and experimental capacities. Appendix B contains the following theoretical computations presented in this chapter: effective prestress, theoretical cracking capacity for the test specimens in the undamaged and damaged state, theoretical maximum capacity for the test specimens in the undamaged and damaged state, maximum capacities of Test Specimens 3, 4, 5, and 6, support conditions, and midspan deflection at the maximum load for Test Specimen 1. Calculation of Effective Prestress In order to calculate theoretical values of the cracking and ultimate loads for the test specimens, it was necessary to determine the effective prestress for stress-relieved prestressing strands. Effective prestress is defined as the stress remaining in the prestressing steel after all losses have occurred (ACI 1996). The total losses that result from prestressing, f T were found by summing the losses due to elastic shortening of concrete, f ES creep of concrete, f CR shrinkage of concrete, f SH and steel stress relaxation, f RE, as shown in Equation 5-1. RESHCRESTfffff (5-1) 110

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111 Equations 5-2, 5-3, 5-4, and 5-5 were used to determine the losses due to concrete elastic shortening, concrete creep, concrete shrinkage, and steel stress relaxation, respectively (Precast/Prestressed Concrete Institute 1999). csESESnfKf (5-2) where K ES = 1.0 for pretensioned members, n = modular ratio, and f cs = compressive stress in concrete at the level of prestressing steel immediately after prestress transfer (ksi). cdscsCRCRffnKf (5-3) where K CR = 2.0 for normal weight concrete and f cds = stress from all superimposed permanent dead and sustained loads after prestressing (ksi). RHSVEKfpsSHSH106.01102.86 (5-4) where K SH = 1.0 for pretensioned members, E ps = modulus of elasticity of prestressing reinforcement (ksi), V/S = volume to surface ratio (in), and RH = average ambient relative humidity (%). CfffJKfSHCRESRERE (5-5) where K RE = 185 ksi, J = 0.14 for 250 ksi grade stress-relieved strand, and C = variable that is a function of the initial stress in the prestressing strands, f pi and of the ultimate stress in the prestressing strands, f pu Since the majority of prestress losses occurred soon after transfer of prestress force to the girder, the section properties were determined for the girder only (Precast/Prestressed Concrete Institute 1999). The gross concrete area of the girder, A g the gross moment of inertia about the major axis for the girder, I g and the neutral axis

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112 were taken for the girder only as shown in Figure 5-1a. The eccentricity of the strands at midspan, e, was taken as 11.83 inches representing the vertical difference between the centroid of the harped strands and the neutral axis of the cross-section. To determine the losses due to elastic shortening and creep, it was necessary to calculate the compressive stress in the concrete at the level of the prestressing steel immediately after prestress transfer with the effects for harped strands, f cs as shown in Equation 5-6 (Precast/Prestressed Concrete Institute 1999). gDgicsicsIeMreAPKf221 (5-6) where K csi = 0.9 for pretensioned members, P i = initial prestress force after anchorage seating losses (kip), r = radius of gyration of girder about the major axis (in), M D = moment due to dead load (kip-in). The initial prestress force after anchorage seating losses, P i was calculated by multiplying the area of prestressing by the initial steel stress, f pi The initial stress of the prestressing steel after anchorage seating losses, f pi was taken as 0.70 of the ultimate strength of the prestressing steel, f pu which was 250 ksi (Precast/Prestressed Concrete Institute 1999, ACI 1996). The area of prestressing was taken as 2.376 in 2 representing 22 strands of 7/16 diameter wire. The modulus of elasticity of the prestressing reinforcement was taken as 27500 ksi and the modulus of elasticity of the concrete, E c was calculated using Equation 5-7 with a concrete compressive strength, f c of the girder of 4500 psi. ccfE'57000 (5-7) where f c is in psi.

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113 The radius of gyration, r, was calculated by Equation 5-8 using the area and moment of inertia corresponding to the gross cross-sectional area of the girder as shown in Figure 5-1a. ggAIr (5-8) The moment due to dead load, M D was calculated using Equation 5-9 with an unsupported length of 40-0 and distributed load, w D as shown in Figure 5-1a. 82wlM (5-9) The stress due to all superimposed permanent dead and sustained loads after prestressing or in this case from the slab, f cds was calculated by Equation 5-10. gsdcdsIeMf (5-10) The moment, M sd was calculated using Equation 5-9 with the distributed load as shown in Figure 5-1b and an unsupported length of 40-0. The volume to surface ratio in Equation 5-4 was calculated by taking the volume of the girder divided by its surface area giving 109.45. The average ambient relative humidity, RH, in Equation 5-4 was taken as 75% (Precast/Prestressed Concrete Institute 1999). The variable C in Equation 5-5 was calculated using Equation 5-11 with the ratio of initial to ultimate strength of prestressing steel, f pi /f pu being 0.70 as assumed above to solve for P i in Equation 5-6. 7.091pupiffC (5-11)

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114 for 0.70 f pi /f pu 0.75. The effective prestress, f pe was calculated as shown in Equation 5-12 by taking the initial stress of the steel, f pi minus the sum of all of the losses from Equation 5-1. Tpipefff (5-12) The resulting effective prestress, f pe was 120 ksi. Theoretical Cracking Capacity for Test Specimens in Undamaged and Damaged State without Repair To calculate the theoretical cracking moment capacity of the test specimens in the undamaged state Equation 5-13 was utilized. DcbgeTgcrMfrecAPyIM'5.712 (5-13) where y T = distance from the centroid of the section to the extreme tensile fibers of the concrete (in) and was taken as 27.3, P e = effective prestress force (kips), c b = distance from the bottom of the section to the neutral axis(in) and equals 15.83 (see Figure 5-1a), and = 1.0 for normal weight concrete. All of the terms inside of the brackets in Equation 5-13 were representative of the girder only and the terms outside of the brackets were representative of the girder with the slab. Therefore, the gross area of concrete, location of the neutral axis, and the input values for the radius of gyration using Equation 5-8 were taken from Figure 5-1a. The effective prestress force, P e was calculated by multiplying the area of prestress, A ps by the effective prestress, f pe The moment due to dead load of the girder with the slab, M D was calculated using Equation 5-9 with an unsupported length of 40-0 and distributed load, w D as shown in Figure 5-1c. Using these input values for Equation 5-13 resulted in a theoretical cracking moment of 1034.73 kip-ft for the live load of the test specimen in the

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115 undamaged state. This theoretical cracking moment for the undamaged specimen corresponds to a cracking load equal to 137.96 kips (see Figure 3-5). The theoretical cracking load for the girders in the damaged state can also be calculated using Equation 5-13. This equation is for an undamaged section and is not directly applicable to a damaged section. However, if Equation 5-13 was used for a damaged section, the calculated cracking moment capacity would be less accurate as a result of the simulated damage. To calculate the damaged theoretical cracking moment capacity the above procedure was followed except that the effective prestress force, P e was calculated by taking the effective prestress, f pe multiplied by the reduced area of prestressing reinforcement, A ps The reduced area of prestress was taken as 1.944 in 2 representing 18 strands or 22 strands minus the four removed strands each being 7/16 in diameter. Using these values as input in Equation 5-13 resulted in a cracking moment capacity of 867.84 kip-ft for the live load of the specimen in the damaged state. This theoretical cracking moment for the damaged specimen corresponds to a cracking load equal to 115.71 kips (see Figure 3-5). Theoretical Ultimate Capacity for Test Specimens in Undamaged and Damaged State without Repair The guidelines suggested by ACI Committee 440 (2002) were used to calculate the theoretical ultimate moment capacity of the test specimens in the undamaged state. ACI guidelines suggest first setting the compressive strain at the top of the section equal to 0.003 in/in to solve through iterations the location of the neutral axis. The location of the neutral axis is then used to solve for the capacities of the steel and concrete. The girder with the slab was used to determine the ultimate capacity for the test specimens in the undamaged and damaged state (see Figure 5-1c). The depth to the individual steel layers

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116 from the top of the section and corresponding area of prestressing steel were taken as shown in Figure 5-2a. The force contribution from the concrete was found by multiplying the stress in the concrete at a given level by the width of the section at that level only for the portion of the section above the neutral axis. Linear interpolation between the strain at the top of the section equal to 0.003 in/in and the strain at the neutral axis location equal to zero was used to determine the strain in the concrete at that given level. The moment contribution of the concrete was found by integrating, from the neutral axis to the top of the section, the force at each level by the corresponding distance to the elastic centroid of the section. The force contribution from the steel was found by integrating over the three layers of steel the area of prestressing at each level times the stress in the prestressing steel at that corresponding level. The Ramberg-Osgood function shown as Equation 5-14 was used for determining the stress in the prestressing steel (Collins and Mitchell 1991). CCpspsBAAEf111 (5-14) The Ramberg-Osgood coefficients, A, B, and C, were found from direct tensile tests of the removed strands of Test Specimen 2, and determined to be 0.1091, 115.09, and 10, respectively (see Figure 3-4). The strain in any given layer of prestressing steel was determined by linear extrapolation with the strain at the top of the section equal to 0.003 in/in and the strain at the neutral axis location equal to zero. The strain input into Equation 5-14 was the sum of the strain in the layer of prestressing from linear extrapolation and the effective prestress strain. The moment contribution of the

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117 prestressing steel was found by summing the force at the level of prestressing times the corresponding distance to the elastic centroid of the section for each layer of prestressing. The total load capacity was determined by taking the force contribution of the prestressing steel minus the force contribution of the concrete. The total moment capacity was determined by summing the moment contributions of the concrete and prestressing steel. Using the above procedure, the total ultimate moment capacity was determined to be 2034.02 kip-ft for the specimen in the undamaged state. The ultimate moment due to live load was determined by taking the resulting ultimate moment capacity minus the dead load moment. The resulting live load ultimate moment capacity was 1897.15 kip-ft. This theoretical ultimate moment for the undamaged specimen corresponds to an ultimate load equal to 252.95 kips (see Figure 3-5). The theoretical ultimate load for the girders in the damaged state can also be calculated by the same procedure. This procedure is for an undamaged section and is not directly applicable to a damaged section. However, if the procedure were used for a damaged section, the maximum moment capacity would be less accurate as a result of the simulated damage. To calculate the damaged theoretical ultimate moment capacity, the above procedure was followed except that the area of steel was reduced for the bottom two layers of prestressing steel. The depth to the individual steel layers from the top of the section and corresponding area of prestressing steel were taken as shown in Figure 5-2b. Using the above procedure, the total ultimate moment capacity was determined to be 1679.59 kip-ft for the specimen in the damaged state. The ultimate moment due to live load was determined by taking the resulting ultimate moment capacity minus the dead load moment. The resulting live load ultimate capacity was 1542.72 kip-ft. This

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118 theoretical ultimate moment for the damaged specimen corresponds to an ultimate load equal to 205.7 kips (see Figure 3-5). Theoretical Ultimate Capacities of Test Specimens 3, 4, 5, and 6 The guidelines suggested by ACI Committee 440 (2002) were used to calculate the theoretical ultimate moment capacity of the test specimens. ACI guidelines suggest first setting the tensile strain at the extreme bottom of the section equal to ultimate strain of the laminate and then solve through iterations the location of the neutral axis. The location of the neutral axis was then used to solve for the capacities of the steel, concrete, and FRP. No factors of safety were used for the analysis, such as the environmental reduction factor, C e to be applied to the ultimate tensile strength or the bond dependent coefficient, to be applied to the rupture strain suggested by the ACI guidelines. The cross-section properties at midspan for the girder with the slab were utilized for the analysis of all of the repaired test specimens (see Figure 5-1c). The depth to the individual steel layers from the top of the section and corresponding area of prestressing steel were taken as shown in Figure 5-2a. The force and moment contribution from the concrete was determined as described in the previous section except that the limits for linear interpolation of the concrete stress have changed. Linear interpolation between the strain at the bottom of the section equal to the rupture strain of the laminate and the strain at the neutral axis location equal to zero was used to determine the stress in the concrete at any level. The force and moment contribution from the steel was determined as described in the previous section except that the limits for linear extrapolation of the steel strain have changed. The strain in any given layer of prestressing steel was determined by linear interpolation between the strain

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119 at the bottom of the section equal to the rupture strain of the laminate and the strain at the neutral axis location equal to zero. The strain input into Equation 5-14 was the sum of the strain in the layer of prestressing from linear interpolation and the effective prestress strain. The force contribution of the FRP was determined by multiplying the cross-sectional area of the FRP by the stress in the FRP. The stress in the FRP was found by multiplying the rupture strain or elongation at break strain by the tensile modulus of elasticity of the FRP. The moment contribution of the FRP was determined by multiplying the force contribution of the FRP by the distance of the FRP to the elastic centroid of the section. The total load capacity of the section was determined by taking the force contribution of the concrete minus the force contribution of the prestressing steel and minus the force contribution of the FRP. The total moment capacity was determined by summing the moment contributions of the concrete, prestressing steel, and FRP. Using the above procedure, the total ultimate theoretical moment capacity for Test Specimens 3, 4, 5, and 6 were determined to be 2914.69 kip-ft, 2326.74 kip-ft, 2206.13 kip-ft (2412.58 kip-ft actual), and 2140.13 kip-ft, respectively. The ultimate theoretical moment capacity due to live load was determined by taking the resulting ultimate moment capacity minus the corresponding dead load moment. The resulting theoretical live load ultimate capacities for Test Specimens 3, 4, 5, and 6 are 2777.82 kip-ft, 2182.94 kip-ft, 2069.26 kip-ft (2275.71 kip-ft actual), and 2003.26 kip-ft, respectively. These theoretical ultimate moment capacities for the repaired specimen correspond to ultimate loads equal to 370.38 kips, 281.67 kips, 275.90 kips (303.43 kips actual), and 267.1 kips, respectively (see Figure 3-5).

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120 The disadvantages to using this procedure were that the FRP thickness on the bottom flange could not be varied as in the case of Test Specimen 5. Also, FRP on the sides of the specimens could not be included as additional flexural capacity only as shear capacity. Support Conditions To determine the behavior of the support conditions during the experimental tests, analytical midspan (maximum) deflections were determined for the undamaged girder for two cases. Case 1 assumes that the girder is a simple beam with two equal concentrated loads symmetrically placed and Case 2 assumes that the girder is fixed at both ends with two loads symmetrically placed. Equations 5-15 and 5-16 represent the deflection equations for Case 1, pinned supports, and Case 2, fixed supports, respectively (AISC 2001). 22max4324)(alEIPacenterat (5-15) for pinned supports, where P = load at one load cell, l = unsupported length and a = distance from support to load point. bxaxalEIlxPbcenterat3362)(322max (5-16) for fixed supports, where b = distance to the load point, x = distance to calculated deflection. The concrete compressive strength, f c is taken as 4500 psi and used in Equation 5-6 to determine the modulus of elasticity of the concrete, E c The unsupported length, l, is taken as 40-0 representing the distance between the supports. The girder with the slab was used to determine the support conditions for the test specimen in the undamaged

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121 state (see Figure 5-1c). The distance to the load point, a, was taken as 15-0 representing the distance between one support and the nearest load point during the experiment. The distance to the load point, b, was taken as 25-0 representing the distance between the load point and the support at the other end of the beam during the experiment. The distance to the deflection, x, was taken as 20-0 representing the distance from one support to where the deflection was being calculated on the beam. The experimental load and midspan deflection chosen for analytical comparison was 98.4 kips and 0.349 where the load is the sum of the two load points on the beam (see Figure 3-5). The experimental load, equal to 98.4 kips, was chosen for the support condition evaluation since it was less than the experimental cracking load for Test Specimen 1 (131.3 kips) as well as the analytical cracking load equal to 132.69 kips. The load, P, to be input into Equations 5-15 and 5-16 was 49.2 kips representing the measured load from one load cell. The analytical midspan deflections determined for the pinned supported case and fixed supported case using a load, P, of 49.2 kips were 0.328 and 0.076, respectively. Therefore, the analytical and experimental results for the deflection at midspan of Test Specimen 1 confirm that the case with pinned supports provides a verification that the supports for the experiments behaved as pinned supports as assumed. Theoretical Midspan Deflection at the Theoretical Ultimate Capacity of the Girder in the Undamaged State In order to normalize the experimental data for comparison it was necessary to determine the ultimate theoretical deflection at midspan. The ultimate deflection at midspan was calculated using Equation 5-15 with the assumption that the supports behave as pinned supports during testing. The maximum load was 135.6 kips which represents the theoretical load at one load point as solved previously for Test Specimen 1.

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122 The area of prestressing, A ps was taken as 2.376 in 2 representing 22 strands each with a diameter of 7/16. The gross moment of inertia can not be used to determine the midspan deflection at the maximum load. For Equation 5-15, the cracked moment of inertia must be used to obtain a deflection representative of the state of the girder at maximum load. The cracked moment of inertia was determined using Equation 5-17 (Precast/Prestressed Concrete Institute 1999). pppscrndnAI6.112 (5-17) where d p = depth to the prestressing strands and p = ratio of prestressed reinforcement. The depth to the prestressing strands, d p was taken as 44 representing the distance from the top of the section to the middle layer of prestressing strands (see Figure 5-2a). The ratio of prestressed reinforcement, p was calculated using Equation 5-18 with the width of the section, b, taken as 24 representing the width of the section in compression. ppspdbA (5-18) The resulting theoretical midspan deflection at ultimate was 5.67. Comparison of Theoretical and Experimental Capacities and Deflections Tables 5-1 and 5-2 provide a comparison between the live load theoretical (T) and experimental (E) capacities at cracking load and at maximum load, respectively. The percent differences, shown in Tables 5-1 and 5-2, were calculated by dividing the difference between the two quantities by the calculated theroetical capacity. Table 5-1 presents the theoretical cracking capacities for Test Specimens 1 and 2 calculated with the assumption of fully composite action existing between the slab and girder. Test Specimen 1 had a theoretical cracking load of 138 kips and an experimental

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123 cracking load of 131 kips while Test Specimen 2 had a theoretical cracking load of 116 kips and an experimental cracking load of 112 kips as shown in Table 5-1. Therefore, a 16.1% reduction in the theoretical cracking load capacity was estimated by the severing of four prestressing strands. Therefore, for Test Specimen 1 the experimental and theoretical capacities at the cracking load for Test Specimen 1 differed by 4.9%. Test Specimen 2 had an experimental cracking load of 112 kips and had a theoretical cracking load of 116 kips as shown in Table 5-1. Therefore, for Test Specimen 2 the experimental and theoretical capacities at the cracking load differed by 3.3%. The calculated theoretical cracking capacity was slightly higher than the observed experimental cracking capacity for Test Specimens 1 and 2. The calculated theoretical cracking capacities were based on the previously determined effective prestress force in the steel of 120 ksi. Although the theoretical cracking capacity was higher than the experimental cracking capacity, the difference can be attributed to several factors such as the assumption that the initial prestressing force should be 70% of the ultimate and the camber of the girders increasing over time due to being out of service without an applied dead load. Test Specimen 1 had a theoretical maximum load of 253 kips and an experimental maximum load of 268 kips while Test Specimen 2 had a theoretical maximum load of 206 kips and an experimental maximum load of 210 kips as shown in Table 5-2. Therefore, an 18.7% reduction in maximum load capacity was estimated by the severing of four prestressing strands, the experimental and theoretical capacities at the maximum load for Test Specimen 1 differed by 6.1%, and the experimental and theoretical capacities at the maximum load for Test Specimen 2 differed by 2.1%. Test Specimen 3 had an experimental maximum load of 244 kips and had a theoretical maximum load of

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124 371 kips as shown in Table 5-2. Therefore, the experimental and theoretical capacities at the maximum load differed by 34.5%. Test Specimen 4 had an experimental maximum load of 202 kips and had a theoretical maximum load of 282 kips as shown in Table 5-2. Therefore, the experimental and theoretical capacities at the maximum load differed by 28.5%. The theoretical capacities presented in Table 5-2 for Test Specimen 5 were the result of a constant FRP laminate thickness of 0.5 with a fiber length of approximately 1.25 applied only to the tension face of the girder. Test Specimen 5 had an experimental maximum load of 255 kips and had a theoretical maximum load of 304 kips as shown in Table 5-2. Therefore, the experimental and theoretical capacities at the maximum load differed by 16.2%. The unrepaired theoretical capacities were consistent with the experimental findings whereas the repaired theoretical capacities were much larger than the observed experimental capacities, which could be due to premature failure modes for Test Specimens 3 and 4, imperfect application of the FRP systems, unconservative repair guidelines set by ACI 440, or FRP material flaws. The theoretical midspan deflection at ultimate for Test Specimen 1 was calculated, with the assumption of pinned supports, as 5.67 and the measured midspan deflection at ultimate load of Test Specimen 1 was 5.95. Therefore, at ultimate load the experimental and theoretical midspan deflection differed by only 6%. The observed midspan deflection at ultimate was slightly higher than the theoretical deflection possibly due to shear deformations being neglected in the midspan deflection calculation with the girder having a shear span to depth ratio of 3.75. Figure 5-3 shows the moment deflection relationship that was normalized to the theoretical live load moment and deflection values at ultimate for Test Specimen 1. The

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125 data shown for Test Specimen 3 only contains values for the last series of loading up to failure. Figure 5-3 shows consistent calculations of the theoretical deflection and capacity of Test Specimen 1 and substantial reductions in deflection and ductility of the girders repaired with FRP systems. Computer Program Implementation for all Test Specimens A computer program was implemented to calculate theoretical moment curvature diagrams and ultimate capacities of the six test specimens for comparison with experimental values (Fung 2002, Consolazio et al. in press). The computer program evaluates capacities based on section properties. The cross-section properties at midspan were input into the computer program for all test specimens. The major advantage of the program was the evaluation of a cross-section including the effects of fiber reinforced polymer systems with any dimensions. The input dimensions are based on a coordinate system for the cross-section contour, steel locations, FRP locations and properties, the concrete compressive strength of the girder, f c and the effective prestress of the prestressing strands, f pe The required steel properties were the yield strength of the mild steel, tensile modulus of the prestressing steel, area of steel, type of steel, and location of the mild steel and/or prestressing steel. Some of the disadvantages of the program were that the cross-section could not be entered as two parts with different concrete compressive strengths as existed in the test specimens, and that all of the FRP failure modes such as bond failures were not considered. By taking the experimental strain profiles and solving for the location of the neutral axis at a specific load, the experimental curvature was determined for that specific load. The calculated experimental curvature has a corresponding load and moment which were

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126 used for comparing the theoretical moment curvature values from the computer program output. Figure 5-4 shows the theoretical moment curvature relations for Test Specimens 1, 2, 3, 4, 5, and 6 and the experimental moment curvature relations for Test Specimens 1, 2, 3, 4, and 5 with T and E denoting the theoretical and experimental relationships, respectively. Figure 5-4 shows three separate theoretical moment curvature relations for Test Specimen 5. The first two series 5a and 5b represent the repair with a fiber length of approximately 2 with a thickness of 0.35 and the repair with a fiber length of approximately 1.25 with a thickness of 0.50. The last series of theoretical moment curvature points, 5c, was generated with the average thickness of the FRP applied at the midspan which was measured from samples taken from the girder after the load test. This program generated the first point at a value of negative curvature. Therefore, the theoretical values shown in Figure 5-4 were offset by the initial negative curvature of the evaluated cross-section such that all of the moment curvature relationships start at zero curvature and zero moment. The theoretical moment curvature points in Figure 5-4 were fairly consistent with the experimental data in the elastic range but vary greatly in the inelastic range. The theoretical ultimate capacities computed by this program are shown in Table 5-3 with E in Column 2 denoting experimental data. The computer program produced consistent moment capacities for the unrepaired specimens with a 1.78% difference for Test Specimen 1 and a 3.74% difference for Test Specimen 2. Generally the computer program produced moment capacities for the repaired specimens higher than the observed experimental capacities. For Test Specimens 3 and 4, the experimental and theoretical maximum moment capacities differed by 32.6% and 30%, respectively. This significant difference can be attributed to the specimens failing prematurely and not

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127 at full capacity. For Test Specimen 5, three analyses were performed (5a, 5b, and 5c) in which the 5c analysis represents the actual repair that was applied to the girder. The experimental and theoretical maximum moment capacities for the 5c analysis differed only by 6.1%.

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Table 5-1: Comparison of Experimental and Theoretical Capacities of Test Specimens 1, 2, 3, 4, 5,and 6 at Cracking Test SpecimenPcr (kip) EMcr (kip-ft) EPcr (kip) TMcr (kip-ft) T% difference of Mcr T of test specimen 1 and T per test specimen% difference of Mcr T and E per test specimen1131.3984.4138.01034.70.05.12111.9839.4115.7867.816.13.4 Table 5-2: Comparison of Experimental and Theoretical Capacities of Test Specimens 1, 2, 3, 4, 5 and 6 at Maximum Load Test SpecimenPmax (kip) EMmax (kip-ft) EPmax (kip) TMmax (kip-ft) T% difference of MmaxT of test specimen 1 and T per test specimen% difference of Mmax T and E per test specimen1268.32012.1253.01897.20.05.72210.11575.9205.71542.718.72.13243.91829.1371.02782.646.752.14201.61562.4282.02185.415.239.95254.81910.8304.12280.620.219.46NANA268.12011.16.0NA 128

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129 Table 5-3: Theoretical Moment Capacity at Ultimate Generated by Computer Program for all Test Specimens Test SpecimenMtest (kip-ft) EMtest (kip-ft)% difference of Mtest experimental per test specimen12012.11976.91.821575.91637.03.731829.12713.932.641562.42224.829.85a1910.82395.620.25b1910.82652.028.05c1910.82035.46.160.02054.3NA

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130 (a) (b) (c) Figure 5-1: Geometric Properties of the Cross-Section a) Girder without the Slab; b) Slab; and c) Girder with the Slab

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131 (a) (b) Figure 5-2: Steel Areas and Location for a) an Undamaged Section; and b) a Damaged Section

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132 Figure 5-3: Normalized Moment versus Deflection for Test Specimens 1, 2, 3, 4, and 5 Figure 5-4: Moment Curvature Comparison of Theoretical Data from Computer Program and Experimental Data

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CHAPTER 6 DESIGN RECOMMENDATIONS AND PRELIMINARY DESIGN SPECIFICATIONS All of the repair design procedures for the FRP strengthening systems of the damaged girders were reasonable and followed typical prestressed and FRP behavior. Each repair design for the damaged girder using FRP used different factors of safety based on engineering judgement and not guideline procedures. All of the design procedures consisted of determining an ultimate capacity, which was not based on a critical limit state or critical failure mode that would be attained if perfect application and bonding was ensured. The biggest issue with repair design procedures using FRP strengthening systems is that code standards do not exist, only guidelines. Guidelines have been established for FRP repair of reinforced concrete structures but only a limited part of these guidelines address or even relate to FRP repair of prestressed concrete girder systems (ACI 2002). For the computation of the ultimate capacity of a reinforced concrete member flexurally strengthened with a FRP system, ACI guidelines suggest setting the tensile strain at the extreme bottom of the section equal to ultimate strain of the laminate. The neutral axis is then determined through iterations with the limit on the extreme tensile strain equal to the laminate rupture strain. The location of the neutral axis can then used to solve for the contributions of the steel, concrete, and FRP. ACI suggests the use of an environmental reduction factor, C e to be applied to the ultimate tensile strength of the laminate, which is a function of the expected exposure condition, fiber type, and resin type as shown in Table 6-1. The purpose of the environmental reduction factor, C e is to 133

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134 take into the account how the mechanical properties of an FRP system are affected by long-term exposure to the environment. ACI also suggests the use of a factor, m to be applied to the rupture strain of the laminate that is computed as shown in Equation 6-1 (ACI 2002). 000,000,1tnEfor90.0tnE000,500601000,000,1tnEfor90.0000,000,2tnE1601fffffufffffum (6-1) where fu = rupture strain of FRP reinforcement, n = number of plies of FRP reinforcement, E f = tensile modulus of elasticity of the FRP, and t f = nominal thickness of one ply of FRP reinforcement. The m factor limits the strain developed in the laminate in order to prevent debonding of the FRP laminate. ACI (2002) also provides the following guidance for the flexural capacity calculation of a prestressed concrete member when the member is strengthened with longitudinal FRP reinforcment to its tension face, In the case of prestressed concrete members, strain compatibility, with respect to the state of strain in the stressed member, should be used to evaluate the FRP contribution. The setting of standards for FRP repairs to reinforced and prestressed concrete would reduce the amount of engineering judgment used in current designs, and therefore would reduce errors and ensure that a proper repair was designed. Although the repair of the test specimens focused mainly on flexural strengthening due to the damage imparted to only the flexural reinforcement, shear capacity must also be an issue in the repair design. Shear resistance is provided through the internal stirrups of prestressed concrete girders. Even though the damage procedure did not involve any

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135 damage to the internal shear stirrups, the shear capacity was reduced by 14.7% experimentally and 6.2% theoretically when compared to the undamaged shear capacity. This reduction in shear capacity can be attributed to the removal and replacement of the 5-0 section centered about midspan. The damage process interrupted the original shear load path. In the undamaged specimen, the load moves through a continuous section without initial cracks. In the damaged specimens, the load moves through a discontinuous section that results from having adjoining concrete of differing properties. To regain the shear capacity through an FRP strengthening system, it is necessary to provide a load path from the FRP being applied primarily for flexure, back to the internal shear stirrups, which were designed for the undamaged shear capacity. A way to ensure that an adequate shear load path will be restored to a flexurally damaged member, is to require that FRP material be applied to the girders for shear reinforcement. For example, if the structural capacity being restored is due to flexural inadequacies, then a cross-sectional area equal to a certain percentage of the applied flexural reinforcement needs to be applied to ensure adequate shear capacity. Based on the findings of this experimental testing program, preliminary design specifications for prestressed concrete girder repairs with FRP materials are as follows: All concrete surfaces should be cleaned and sandblasted prior to FRP application. Cracks, voids, and discontinuities should be repaired to ensure that there will be a good bond of the FRP materials. FRP materials used for flexural reinforcement on the tension face of a structural member should extend, at a minimum, 0.80 times the unsupported length For flexurally repaired structural members, an area equal to a percentage of flexural reinforcement should be applied to the sides of a member for shear strength. This percentage needs to be determined through further investigation;

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136 Stirrups should be used at the termination points of FRP flexural reinforcement for proper anchorage and to avoid the failure modes due to concrete cover separation and FRP debonding; The stirrups should be made of bi-directional fabric with fibers oriented in the direction of and perpendicular to the longitudinal axis of the structural member;

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137 Table 6-1:Environmental Reduction Factor for Various FRP Systems and Exposure Conditions (ACI 2002) Exposure ConditionsFiber and resin typeEnvironmental reduction factor, CeCarbon/epoxy0.95Glass/epoxy0.75Aramid/epoxy0.85Carbon/epoxy0.85Glass/epoxy0.65Aramid/epoxy0.75Carbon/epoxy0.85Glass/epoxy0.50Aramid/epoxy0.70Interior exposureExterior exposure (bridges, piers, and unenclosed parking garages)Aggressive environment (chemical plants and waste water treatment plants)

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CHAPTER 7 SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS FOR FUTURE RESEARCH A short term way to avert collisions due to over-height vehicles would be to regulate that all bridges including underpasses to the Interstate System meet the current minimum desired height requirements. This would mean that pre-existing bridges would have to be replaced or lifted, or underlying roadways would have to be lowered requiring a tremendous amount of work and funding. Also this would mean that in a couple of years vehicle heights would increase and bridges would again be subject to impacts due to over-height vehicles resulting in a continuous cycle of increased bridge heights followed by increased vehicle heights. Until drastic measures are taken to ensure that collisions due to over-height vehicles will not occur, it is necessary to investigate emergency repair methods, especially of prestressed concrete bridges. The repair of prestressed concrete structures with FRP systems needs additional research for the following reasons: to determine that all failure modes have been observed, to determine solutions to undesirable failure modes, to determine if the necessary capacity of a damaged member can be regained through FRP strengthening, and to determine standards for the repair designs of FRP strengthened systems of prestressed concrete structures. To minimize over-height vehicle collisions some preventative measures could be used such as additional warning signs, more aggressive enforcement, higher fines for violators, placement of video cameras on previously impacted bridges, or over-height vehicle detection systems with communication devices. 138

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139 Summary Full-scale girder tests of six type II AASHTO girders were performed. The six tests represented an undamaged control specimen, a control specimen with simulated damage, and the four specimens with simulated damage that were then repaired with different fiber reinforced polymer systems. The FRP systems varied in material types, laminate properties, application procedures, reinforcing schemes, and repair design procedures and assumptions. The specimens were tested to failure to determine moment and shear capacities, as well as deformation and ductility behavior of the undamaged, damaged, and repaired girders. Through experimental and analytical comparisons of the behavior of the repaired girders to the undamaged girder, the specific fiber reinforced polymer systems were evaluated for acceptance to the FDOT Quality Products List. The FRP systems were evaluated structurally on their ability to restore the undamaged shear and moment capacity of the prestressed concrete girder, type of failure mode, cost of repair system including installation and ease of installation. Conclusions Although Test Specimen 3 failed prematurely, it still regained 89.9% of the theoretical capacity and 90.9% of the experimental capacity of Test Specimen 1. The use of stirrups at the ends of the FRP laminate could have provided sufficient resistance to the peeling stresses that led to premature failure. Test Specimen 4, repaired with carbon fibers and a polyurethane adhesive, failed prematurely due to adhesive failure followed by FRP rupture. The test specimen did not regain any of the lost capacity due to the simulated damage. Test Specimen 5 failed due to FRP rupture with a regained theoretical capacity equal to 95% and with a regained experimental capacity equal to 95% of Test Specimen 1. The desired thickness of the spray FRP for the tensile face of the beam was

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140 0.50, although only an average thickness of 0.27 was achieved. If the desired thickness of 0.50 had been achieved, Test Specimen 5 could have, at a minimum, attained the same capacity as the undamaged girder. The midspan deflections at the maximum capacity for all of the repaired test specimens showed significant reductions with a minimum reduction equal to 44% of the theoretical deflection at maximum capacity of Test Specimen 1. The measured strains on the bottom centerline of the girder at maximum capacity for all of the repaired test specimens showed significant reductions with a minimum reduction equal to 49% of the experimental maximum strain at maximum capacity of Test Specimen 1. Future Testing Recommendations for Measurements to Assess Adequacy and Performance of FRP Repair The lack of instrumentation and properly functioning instruments used for the girder tests resulted in minimal conclusions being drawn from the experiments in some cases. By increasing the number and type of instruments used on each specimen, better information can be obtained to help understand what is occurring in the specimens during loading. Additional load cells at the supports are needed to determine if unequal loading is resulting in unequal reactions at the supports. More strain gauges are needed in the compression region to make up for gauges that were located too close to the neutral axis resulting in unusable data. If possible, strain gauges should be placed directly on the prestressing strands to obtain more accurate data than from the crack gauges applied to the concrete surface at the level of prestressing. By instrumenting the prestressing strands directly, the problems encountered with the crack gauges might have been avoided. Instrumenting both the damaged and undamaged prestressing strands would show if the damaged strands should be fully or partially neglected in repairs and if the

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141 damaged strands have similar strains to the surrounding concrete over only a portion of their length due to their loss of development and transfer length. Instrumenting the damaged and undamaged strands would also provide information related to strand slip or strand failure during testing. For future tests, camber measurements before the slab is poured, after the slab is cast, after the concrete is removed (damage), after the strands are severed, after the concrete patch is applied, and after the FRP system is applied will all be valuable for determining stresses in the test specimens.

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APPENDIX A REPAIR DESIGNS This appendix contains the design procedures for the repairs of Test Specimens 3, 4, and 5. RJWatson Repair Design The RJWatson repair design is shown below as provided to the FDOT for installation. CLIENT: Florida DOT SHEET 1 OF 2 PROJECT NAME: Test Program PROJECT #: CALCULATION BY: SEW DATE: 11-15-02 Design Goal: To provide additional flexural capacity to the damaged beams. Existing Beam Properties: b1 86 beam width (on bottom) (in) d3 beam depth (in) ts12 slab thickness (in) dfdts total depth of beam and slab for flexure (in) df48 l43.75 beam span length (ft) f'c4.5 concrete compressive strength (ksi) Tyfo SCH 41S Unidirectional Carbon Composite Properties: Eu1050 0 typical test value for tensile modulus (ksi) CE0.85 environmental reduction factor for carbon ECEEu guaranteed design value for tensile modulus (ksi) E8925 0.006 allowable design strain in composite for flexure 142

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143 fE guaranteed design stress in composite (ksi) f53.6 t0.04 layer thickness of SCH 41S composite (in) Moment capacity enhancement with one layer of composite: The moment capacity enhancement of the beam provided by the composite is: M = A f (jd) 0.9 ACI reduction factor for flexure At b area of one layer of composite (in2) A0.72 jd0.9df assumed moment lever arm (in) jd43.2 MAfjd12 M124.9 additional moment capacity of beam with one layer of composite (k-ft) Required number of layers of composite: The number of layers of composite needed on the beam is based on the additional moment demand required. Mreq20121576 required additional moment capacity (k-ft) layersMreqM number of layers required layersceillayers() round up to closest number of layers layers4 Design Conclusions: Providing 4 layers of the Tyfo SCH 41S System will enhance the flexural capacity of the beam to restore the original design strength. Air Logistics Repair Design The Air Logistics repair design is shown below as provided to the FDOT for installation except that the written calculations have been duplicated in computer format for presentation.

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144 damaged0.021inin control0.018inin damaged0.027inin control.0095inin is for maximum strain at center on bottom at failure and is for maximum strain on bottom at load points 15 feet from supportsExperimental Strainsdamaged5.57incontrol5.58indamaged6.37incontrol6.34in is for deflection at midspan and is for deflection at the load points 15 feet from supportsExperimental DeflectionsPdamagedE105.1kipPcontrolE134.1kipMdamagedE1576kipftMcontrolE2012kipftExperimental CapacitiesPdamagedT108.3kipPcontrolT133.8kipMdamagedT1624kipftMcontrolT2007kipftTheoretical CapacitiesTEST DATA Goal: Strengthen the damaged girder in a manner that results in the girder resembling or exceeding the performance of the control girder Objective: Use CFRP retrofit system to strengthen the damaged girder Structure: AASHTO Type II Girder-Control-Damaged: 4 prestressing strands severed Project: FLDOT AASHTO Girder Test Program Client: Air Logisitcs Corporation-FDOTksi1000psikip1000lbf

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145 0102030 40 1107 5106051061107 Distance (ft)Shear (kips)Vsupport234.67kipVsupportP1lx()l P2lxs()l Msupport0kipftMsupportxlx()l P1P2xs()x x0ftVloadpoint134.1kipVloadpointP1lx()l P2lxs()l Mloadpoint1676.25kipftMloadpointxlx()l P1P2xs()x x15ftVmidspan0kipMmidspan2011.5kipftMmidspanxlx()l P1P2xs()x x20ftP2P1s10ftP1PcontrolEl40ft

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146 01020304005001000150020002500 Distance (ft)Moment (kip-ft) 010203040010203040 girder slab strands rebar damagedAg12in24in()12in6in()6in3in()2.53in3in()6in15in6in6in()2.56in6in()6in18in()Ag657in2Ag4.56ft2

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147 ts24innEsEc n6.41ds12inAstop0.31in2Area of #5 bars at midheight of slab Area of strand of 7/16 inch diameter and 250 ksi steelAs.115in2Transformed Areas As16AsAs2control8AsAs2damaged6AsAs3control8AsAs3damaged6AsAcontrolAs1As2controlAs3control AdamagedAs1As2damagedAs3damaged Concrete AcAgAc4.56ft2VcAg43.75ftVc199.61ft3150lbfft3 WVcW29.94kipwW43.75ft w0.684kipft Mdeadwl28 moment at midspan due to self-weightMdead136.87kipftCalculate Cracked Transformed Section d142inf'c6300psicompression is in the slabd244inEs29000ksid346inEc57000psi.5f'c Ec4524.23ksid48in

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148 ccontrolcXc()0.01in3comparable to data points from tested strain gaugesc7.0462inXc()tscc2 Asposcds2 Ascontrolcd2CASE 1: CONTROLCalculate Neutral Axis for Three Cases for 4 layersAf43.27in2Af4EfEc wfztfnumber of layersz4for 3 layersAf32.45in2Af3EfEc wfztfnumber of layersz3fabric.012inin ffabric120ksiwf12intf0.028inEf11000ksiBottom CarbonAsdamaged13.27in2AsdamagednAs1As2damagedAs3damagedonly 18 strands carrying load for damaged conditionAscontrol16.22in2AscontrolnAs1As2controlAs3controlAspos3.35in2Aspos2Astopn1()Steel

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149 0.01 0.00500.00502040 section strain steel 1 steel 2 steel 3 zero FRPSTRESS DISTRIBUTIONXc()0.01in3crepaired3cc7.0004inXc()tscc2 Asposcds2 Asdamagedcd2Af3cd()for 3 layers of FRPXc()0in3crepaired4cc7.1766inXc()tscc2 Asposcds2 Asdamagedcd2Af4cd()for 4 layers of FRPCASE 3: REPAIREDXc()0.01in3cdamagedcc6.4354inXc()tscc2 Asposcds2 Asdamagedcd2CASE 2: DAMAGED

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150 jd39.0141.0143.01 injdd1a2 d2a2 d3a2 a0.85ccontrolCONTROL CALCULATE MOMENT CAPACITY ff1f2f3 f3fuf1224.33ksif1d1ccontrolfud3ccontrol f2237.16ksif2d2ccontrolfud3ccontrol CONTROL f2237.12ksif2d2crepaired4fud3crepaired4 REPAIRED 4 LAYERSFind Steel Stresses by Similar TrianglesAt failure assume 1) concrete is at limit cu = 0.003 2) bottom steel is at ultimate limit fu = 250 ksi (3 layers of steel) 3) other two layers of steel at reduced stressinitial tendon stressfi180.62ksifi0.85fyfy212.5ksify0.85fufu250ksiASSUMPTIONS

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151 Mr42162.09kipftMr402iAdamagedifijdi ztfwfffabricjdfjd38.9540.9542.95 injdd1a2 d2a2 d3a2 a0.85crepaired4z4with 4 layersby setting the ultimate strain for the FRP fabric equal to the ultimate strain found experimentally, a strength reduction of = 0.8 has to be applied for capacity calculations 0.80.792controlfabric jdf45inREPAIRED Md1691.14kipftMd02iAdamagedifijdijd39.2641.2643.26 injdd1a2 d2a2 d3a2 a0.85cdamagedDAMAGED Mc2072.98kipftMc02iAcontrolifijdi

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152 with 3 layersz3a0.85crepaired3jdd1a2 d2a2 d3a2 jd39.0241.0243.02 inMr302iAdamagedifijdi ztfwfffabricjdfMr32044.2kipftM01676.252011.51676.250 kipftMdamagedMdMdMdMdMd Mrepair4Mr4Mr4Mr4Mr4Mr4 Mrepair3Mr3Mr3Mr3Mr3Mr3 051015202530354005001000150020002500 P = 134 kips damaged 4 layers 3 layersDistance (ft)Moment (kip-ft)

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153 P1140kipP2P1x20ftMmidspanxlx()l P1P2xs()x x0ftMsupportxlx()l P1P2xs()x x15ftMlpxlx()l P1P2xs()x M140MsupportMlpMmidspanMlpMsupport xx0ft15ft20ft25ft40ft 051015202530354005001000150020002500 damaged P = 140 kips 4 layers 3 layersDistance (ft)Moment (kip-ft)

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154 051015202530354045024 girder supports fiber 1 fiber 2 fiber 3 fiber 4Alternatively, all 4 layers could be carried over full lengthlayer416ftlayer324ftto be centered at midspanlayer232ftlayer140ftProposed Lengths The numbers inside the first parenthesis to be subtracted above represents the theoretical cutoff and the numbers inside the second parenthesis to be added above represents the providing of Le beyond the theoretical cut off.layer47ftlayer4l220ft()23.5ft()layer312ftlayer3l218ft()24ft()layer218ftlayer2l216ft()25ft()layer127ftlayer1l214ft()27.5ft()Minimum Theoretical Strip Lengths UF Repair Design-As Designed The UF repair design is shown below as it was initially designed and provided to the FDOT

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155 d146(in)distance of first layer of steel from the top of sectiond244(in)distance of second layer of steel from the top of sectiond342(in)distance of third layer of steel from the top of sectionEp27500(ksi)Modulus of elasticity of prestressing steel typical for 250 ksi steelfpe120(ksi)Inputted value from calculation of prestress lossesFRP Ef1711(ksi)Modulus of FRP for spray-up from Boyd's thesis of 2 in fiber length fu0.0132Material failure strain for spray-up from Boyd's thesis of 2 in fiber lengthfiber0.0132ffiberfiberEfFfiberwidthtfffiberMfiberFfiberhtf2 RefCentY h48(in)height of sectionwidth18(in)width of flange that the FRP will be applied to, to assist in carrying momentf'c4.5(ksi)Input thicknesstf0.35(in)Thickness of FRP to be applied varied until desired moment of 2012 kip-ft is reachedRefCentY1359948widthtfwidth2 tf2657widthtf Elastic CentroidRefCentY20.96(in)STEEL Ap160.108(in2)1st layer of steel from the bottomAp260.108(in2)2nd layer of steel from the bottomAp360.108(in2)3rd layer of steel from the bottom

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156 CONCRETE Concrete compressive strain at the top of the section in terms of strain in FRP being at ultimatetopc()fuchc ccz()topc()c zfcscz()nf12.5 f'c0.8fcf'c1265f'c 1000 f'c2.5 0.8 f'c2.5 0.8 1 kf'c9 0.67peakexplnnf1()1nfk() nfk() fcfc_peaknfpeakfc nf1peakfc nfk f'cadjustf'cfc_peak fcs_actualnfccz()fc nf1ccz()fc nfk f'cadjustfcs_actual

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157 Fptotalc()Fp1c()Fp2c()Fp3c()Fp3c()Ap3fp3cFp2c()Ap2fp2cFp1c()Ap1fp1cp3c()pe3c()p2c()pe2c()p1c()pe1c()3c()d3c()d3cc 2c()d2c()d2cc 1c()d1c()d1cc fc()ifEpc()A1A1Bc()()10 0.1 250Epc()A1A1Bc()()10 0.1 250 pe4.363636103pefpeEp Stress relieved steelB115.09A0.01091PRESTRESSRamberg-Osgood Coefficients obtained from direct tensile tests of strands removed from test specimen 2. Mccc()0czzchRefCentY()[][]bcz()fcscz()dMoment contribution from concrete about the elastic centroidForce contribution from concreteFccc()0czbcz()fcscz()dFor calculation with effects of the 12 in slab (in)bcz()24czc12()if12c12()zc18()if62cz()21[]c18()zc21()if6c21()zc36()if6236cz()[][]c36()zc42()if18c42()zc48()if0otherwise Width of Section: z is the distance from the neutral axis; c is the neutral axis position from the top fiber of concrete

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158 Mp1c()Fp1c()d1hRefCentY()[]Mp2c()Fp2c()d2hRefCentY()[]Mp3c()Fp3c()d3hRefCentY()[]Mptotalc()Mp1c()Mp2c()Mp3c()Ftotalc()Fccc()Fptotalc()FfiberMtotal440c()Mccc()Mptotalc()MfiberccrootFtotalc()c015() cc7.801Neutral axis distance from the top of the sectionFcccc()611.11Mtotal440cc()26473.61(kip-in)Ultimate moment capacity Mtotal440cc()12 2206.13(kip-ft)Ultimate moment capacity UF Repair Design-As Built The UF repair design as presented previous is shown below to demonstrate the capacity of the repair with the as built properties with a fiber length of 1.25 and a thickness of 0.50. h48(in)height of sectionwidth18(in)width of flange that the FRP will be applied to, to assist in carrying momentf'c4.5(ksi)Input thicknesstf0.50(in)RefCentY1359948widthtfwidth2 tf2657widthtf Elastic CentroidRefCentY21.071(in)

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159 (ksi)Modulus of elasticity of prestressing steel typical for 250 ksi steelfpe120(ksi)Inputted value from calculation of prestress lossesFRP Ef1522(ksi)Modulus of FRP for spray-up from Boyd's thesis of 1.25 in fiber length fu0.0143Material failure strain for spray-up from Boyd's thesis of 1.25 in fiber lengthfiber0.0143ffiberfiberEfFfiberwidthtfffiberMfiberFfiberhtf2 RefCentY CONCRETE Concrete compressive strain at the top of the section in terms of strain in FRP being at ultimatetopc()fuchc ccz()topc()c zSTEEL Ap160.108(in2)1st layer of steel from the bottomAp260.108(in2)2nd layer of steel from the bottomAp360.108(in2)3rd layer of steel from the bottomd146(in)distance of first layer of steel from the top of sectiond244(in)distance of second layer of steel from the top of sectiond342(in)distance of third layer of steel from the top of sectionEp27500

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160 fcscz()nf12.5 f'c0.8fcf'c1265f'c 1000 f'c2.5 0.8 f'c2.5 0.8 1 kf'c9 0.67peakexplnnf1()1nfk() nfk() fcf.c_peaknfpeakfc nf1peakfc nfk f'cadjustf'cf.c_peak fcs_actualnfccz()fc nf1ccz()fc nfk f'cadjustfcs_actual Width of Section: z is the distance from the neutral axis; c is the neutra l axis position from the top fiber of concretebcz()24czc12()if12c12()zc18()if62cz()21[]c18()zc21()if6c21()zc36()if6236cz()[][]c36()zc42()if18c42()zc48()if0otherwise (in)For calculation with effects of the 12 in slab

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161 Neutral axis distance from the top of the section cc8.327ccrootFtotalc()c015()Mtotal440c()Mccc()Mptotalc()MfiberFtotalc()Fccc()Fptotalc()FfiberMptotalc()Mp1c()Mp2c()Mp3c()Mp3c()Fp3c()d3hRefCentY()[]Mp2c()Fp2c()d2hRefCentY()[]Mp1c()Fp1c()d1hRefCentY()[]Fptotalc()Fp1c()Fp2c()Fp3c()Fp3c()Ap3fp3cFp2c()Ap2fp2cFp1c()Ap1fp1cp3c()pe3c()p2c()pe2c()p1c()pe1c()3c()d3c()fuhc 2c()d2c()fuhc 1c()d1c()fuhc fcifEpc()A1A1Bc()10 0.1 250Epc()A1A1Bc()10 0.1 250 pe4.363636103pefpeEp B115.09Stress relieved steelRamberg-Osgood Coefficients obtained from direct tensile tests of strands removed from test specimen 2 A0.01091PRESTRESS Moment contribution from concrete about the elastic centroidMccc()0czzchRefCentY()[][]bcz()fcscz()dForce contribution from concreteFccc()0czbcz()fcscz()d

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162 Fcccc()665.312Mtotal440cc()28950.94(kip-in)Ultimate moment capacity Mtotal440cc()12 2412.58(kip-ft)Ultimate moment capacity

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APPENDIX B THEORETICAL EVALUATION TOOLS This appendix contains the MathCAD sheets used for calculating theoretical values that were referenced in Chapter 5. Effective prestress The following demonstrates how the effective prestress of the prestressing strands was calculated. conc0.15kipft3 Vc177120in3VcAglPi415.8kipPifpiApsfpi175ksifpi0.7fpur11.754inrIgAg e11.83ine32inybarprestensionedKci0.9n7.192nEpsEc Ec3.824103ksiEc57000f'c psi0.5ybar20.17inl40ftf'c4500psiAps2.376in2Aps220.108in2Ag369in 2 Ig50979in4Eps27500ksifpu250ksiksi1000psikip1000lbf 163

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164 wdeadWdeadl wdead0.3kipft Mdeadwdeadl28 Mdead60kipftfcdsMdeadeIg fcds0.167ksifcrKcrnfcsfcdsfcr23.882ksiLosses due to creepKsh1.0pretensionedRH75perimeter26in6in15in6in9in18 in72 inperimeter109.456inVtoSAgperimeter VtoS3.371inLosses due to shrinkagefsh8.2106inin KshEps10.06in VtoS 100RH()fsh4.497ksiKre18.5ksiWswconcVcWsw15.375kipwswWswl wsw0.384kipft MDwswl28 MD76.875kipftfcsKciPiAg 1e2r2 MDeIg fcs1.827ksiKes1.0pretensionedfesnfcsKesfes13.143ksiLosses due to elastic shorteningKcr2.0normal weight concreteAdead24in12inVdeadAdeadlVdead80ft3WdeadconcVdeadWdead12kip

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165 J0.14fpifpu 0.7C19fpifpu 0.7 C1Losses due to relaxation of the prestressing strandsfreKreJfesfcrfsh Cfre12.687ksiftotalfesfcrfshfreftotal54.209ksifpefpiftotal fpe120.791ksiEffective prestress Cracking Capacity for Test Specimens in Undamaged and Damaged State The following demonstrates how the theoretical capacity at cracking was calculated for the test specimens in the undamaged and damaged state. YBslab27.3i n YBslab48inYTYTYNAYNA20.7inYNAAybarAslab Aybar13599in3Aybar24in12in()6in12in6in()15in6in3in()19.5in2.53in3in()19in6in15in()28.5in6in6in()39in2.56in6in()40in18in6in45inAslab657in2Aslab24in12in12in6in6in3in2.53in3in()6in15in6in6in2.56in6in()18in6inYBgirder15.83inYBgirder36inYTgirderYTgirder20.17inAgirder369in2given valuesIgirder50979in4ksi1000psikip1000lbf

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166 load at cracking for both load cells Pcr137.96kipPcrMcr15ft 2 Mcr1034.728kipftMcrIslabYBslab PeAgirder 1ecbr2 7.5f'c 1000 kipin2 MdPe285.12kipPefpeApsAps220.108in2fpe120ksiPexp131.2538kipUNDAMAGEDMd136.87kipftMdwl28 w0.68kipft wVslabl 0.15kipft3 VslabAslabll40ftf'c4500Cracking Moment Equation from Nawy Prestressed Concrete text 3rd edition Page 403 Eqn 7.2bnormal wt concrete1.0cb15.83incbYBgirdere11.83ine36in4inYTgirderr11.75inrIgirderAgirder Islab165220.83in4Islab112 24in12in()324in12inYNA6in2112 12in6in()312in6inYNA15in2112 6in3in()36in3inYNA19.5in22136 3in3in()312 3in3inYNA19in2 112 6in15in()36in15inYNA28.5in2112 6in6in()36in6inYNA39in22136 6in6in()312 6in()2YNA40in2 112 18in6in()318in6inYNA45in2

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167 DAMAGEDPexp107.88kipfpe120ksiAps180.108in2PefpeApsPe233.28kipMcrIslabYBslab PeAgirder 1ecbr2 7.5f'c 1000 kipin2 Md Mcr867.84kipf t PcrMcr15ft 2 Pcr115.71kipload at cracking for both load cells Ultimate Capacity for Test Specimens in Undamaged and Damaged State The following demonstrates how the theoretical capacity at ultimate was calculated for the test specimens in the undamaged state. This method can be used to calculate the theoretical capacity at ultimate for test specimens in the undamaged state by reducing the area of the second and third layers of prestressing steel. (in2)3rd layer of steel from the bottomd146(in)distance of first layer of steel from the top of sectiond244(in)distance of second layer of steel from the top of sectiond342(in)distance of third layer of steel from the top of sectionEp27500(ksi)Modulus of elasticity of prestressing steel typical for 250 ksi steelfpe120(ksi)Inputted value from calculation of prestress lossesh48(in)height of sectionf'c4.5(ksi)Elastic CentroidRefCentY13599657 RefCentY20.699(in)STEEL Ap180.108(in2)1st layer of steel from the bottomAp280.108(in2)2nd layer of steel from the bottomAp360.108

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168 CONCRETE topc()0.003Concrete compressive strain at the top of the sectionccz()topc()c zfcscz()nf12.5 f'c0.8fcf'c1265f'c 1000 f'c2.5 0.8 f'c2.5 0.8 1 kf'c9 0.67peakexplnnf1()1nfk() nfk() fcf.c_peaknfpeakfc nf1peakfc nfk f'cadjustf'cf.c_peak fcs_actualnfccz()fc nf1ccz()fc nfk f'cadjustfcs_actual

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169 p3c()pe3c()p2c()pe2c()p1c()pe1c()3c()d3c()d3cc 2c()d2c()d2cc 1c()d1c()d1cc fc()ifEpc()A1A1Bc()()10 0.1 250Epc()A1A1Bc()()10 0.1 250 pe4.363636103pefpeEp Stress relieved steelB115.09A0.01091PRESTRESSRamberg-Osgood Coefficients obtained from direct tensile tests of strands removed from test specimen 2. Mccc()0czzchRefCentY()[][]bcz()fcscz()dMoment contribution from concrete about the elastic centroidForce contribution from concreteFccc()0czbcz()fcscz()dFor calculation with effects of the 12 in slab (in)bcz()24czc12()if12c12()zc18()if62cz()21[]c18()zc21()if6c21()zc36()if6236cz()[][]c36()zc42()if18c42()zc48()if0otherwise Width of Section: z is the distance from the neutral axis; c is the neutral axis position from the top fiber of concrete

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170 Ultimate moment capacity (kip-ft)Mtotal440cc()12 2034.021Ultimate moment capacity (kip-in) Mtotal440cc()24408.253Fcccc()594Neutral axis distance from the top of the section cc7.435ccrootFtotalc()c148()Mtotal440c()Mccc()Mptotalc()Ftotalc()Fccc()Fptotalc()Mptotalc()Mp1c()Mp2c()Mp3c()Mp3c()Fp3c()d3hRefCentY()[]Mp2c()Fp2c()d2hRefCentY()[]Mp1c()Fp1c()d1hRefCentY()[]Fptotalc()Fp1c()Fp2c()Fp3c()Fp3c()Ap3fp3cFp2c()Ap2fp2cFp1c()Ap1fp1c Ultimate Capacity for Repaired Test Specimens 3, 4, 5, and 6 The following demonstrates how the theoretical capacity at ultimate was calculated for Test Specimen 3. This method can be used to calculate the theoretical capacity for Test Specimens 4, 5, and 6 by making the necessary changes of fiber properties, laminate thickness, and laminate width.

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171 distance of first layer of steel from the top of sectiond244(in)distance of second layer of steel from the top of sectiond342(in)distance of third layer of steel from the top of sectionEp27500(ksi)Modulus of elasticity of prestressing steel typical for 250 ksi steelfpe120(ksi)Inputted value from calculation of prestress lossesFRP Ef10500(ksi)Modulus of FRP provided by manufacturer fu0.0121Material failure strain provided by manufacturerfiber0.0121ffiberfiberEfFfiberwidthtfffiberMfiberFfiberhtf2 RefCentY CONCRETE Concrete compressive strain at the top of the section in terms of strain in FRP being at ultimatetopc()fuchc h48(in)height of sectionwidth16(in)width of flange that the FRP will be applied to, to assist in carrying momentf'c4.5(ksi)Input thicknesstf0.16(in)RefCentY1359948widthtfwidth2 tf2657widthtf Elastic CentroidRefCentY20.805(in)STEEL Ap160.108(in2)1st layer of steel from the bottomAp260.108(in2)2nd layer of steel from the bottomAp360.108(in2)3rd layer of steel from the bottomd146(in)

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172 ccz()topc()c zfcscz()nf12.5 f'c0.8fcf'c1265f'c 1000 f'c2.5 0.8 f'c2.5 0.8 1 kf'c9 0.67peakexplnnf1()1nfk() nfk() fcf.c_peaknfpeakfc nf1peakfc nfk f'cadjustf'cf.c_peak fcs_actualnfccz()fc nf1ccz()fc nfk f'cadjustfcs_actual Width of Section: z is the distance from the neutral axis; c is the neutra l axis position from the top fiber of concretebcz()24czc12()if12c12()zc18()if62cz()21[]c18()zc21()if6c21()zc36()if6236cz()[][]c36()zc42()if18c42()zc48()if0otherwise (in)For calculation with effects of the 12 in slab

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173 Neutral axis distance from the top of the section cc9.928ccrootFtotalc()c015()Mtotal440c()Mccc()Mptotalc()MfiberFtotalc()Fccc()Fptotalc()FfiberMptotalc()Mp1c()Mp2c()Mp3c()Mp3c()Fp3c()d3hRefCentY()[]Mp2c()Fp2c()d2hRefCentY()[]Mp1c()Fp1c()d1hRefCentY()[]Fptotalc()Fp1c()Fp2c()Fp3c()Fp3c()Ap3fp3cFp2c()Ap2fp2cFp1c()Ap1fp1cp3c()pe3c()p2c()pe2c()p1c()pe1c()3c()d3c()fuhc 2c()d2c()fuhc 1c()d1c()fuhc fcifEpc()A1A1Bc()10 0.1 250Epc()A1A1Bc()10 0.1 250 pe4.363636103pefpeEp B115.09Stress relieved steelRamberg-Osgood Coefficients obtained from direct tensile tests of strands removed from test specimen 2 A0.01091PRESTRESS Mccc()0czzchRefCentY()[][]bcz()fcscz()dMoment contribution from concrete about the elastic centroidForce contribution from concreteFccc()0czbcz()fcscz()d

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174 Fcccc()793.364Mtotal440cc()34976.31(kip-in)Ultimate moment capacity Mtotal440cc()12 2914.69(kip-ft)Ultimate moment capacity Support Conditions The following demonstrates how the support conditions were determined for Test Specimen 1. 0.0701in 2P2 b2x26EIgl3 3al()3ax()bx[] x20ftb25ftFor fixed supports 0.328in P2 a24EIg 3l24a2a15ftFor pinned supports E3823.68ksiE57000f'c psi0.5f'c4500psil40ftLoad at above given deflection before cracking P98.4153kiptotal centerline deflection before the cracking load exp.34895inIg165220.83in4Worksheet to determine the support conditions based on deflection behavior for the undamaged girderksi1000psikip1000lbf

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175 Deflection at Ultimate for Test Specimen 1 The following demonstrates how the theoretical deflection at ultimate was determined for Test Specimen 1. 5.669inPtheoreta24EcIcr 3l24a2a15ftl40ftIcr26349.392in4IcrnApsdp211.6np pApsbdp b24inAps220.108in2dp44innEsEc Ec3823.676ksiEc57000f'c psi0.5f'c4500psiEs27500ksiPtheoret135.601kipPtheoretMtheoret15ft Mtheoret2034.021kipftWorksheet to determine the theoretical ultimate deflection based on pinned supports and the theoretcial ultimate capacity determined for the undamaged girderksi1000psikip1000lbf

PAGE 190

REFERENCES American Association of State Highway and Transportation Officials (AASHTO) (1990). A policy of geometric design of highways and streets, Washington D.C.: AASHTO. American Concrete Institute (ACI) Committee 440 (2002). Guide for the design and construction of externally bonded FRP systems for strengthening concrete structures. American Concrete Institute, ACI 440.2R-02, Farmington Hills, Michigan. American Concrete Institute (ACI) Committee 440 (1996). State-of-the-art report on fiber reinforced plastic reinforcement for concrete structures. American Concrete Institute, ACI 440R-96, Farmington Hills, Michigan. American Institute of Steel Construction (AISC) (2001). Manual of steel construction, load & resistance factor design, Third Edition, American Institute of Steel Construction, Chicago: AISC. American Society for Testing and Materials (ASTM) (2000). Standard test method for tensile properties of polymer matrix composite materials. D 3039-00, West Conshohocken, Pennsylvania. Arockiasamy, M. (1995). Experimental studies on the feasibility of use of carbon fiber reinforced plastics in repair of concrete bridges. Center for Infrastructure and Constructed Facilities, Department of Ocean Engineering, Florida Atlantic University, Boca Raton, Florida. Arockiasamy, M. and Barbosa, M. (2000). Evaluation of conventional repair techniques for concrete bridges. Center for Infrastructure and Constructed Facilities, Department of Ocean Engineering, Florida Atlantic University, Boca Raton, Florida. Bridge Engineering Software and Technology (BEST) Center (2001). Maryland study, vehicle collisions with highway bridges. Department of Civil and Coastal Engineering, University of Maryland, College Park, Maryland. Boyd, A.J. and Banthia, N. (2001). Strengthening of full-scale bridge channel beams with sprayed FRP. Proceedings of the Ninth International Conference on Structural Faults and Repair, edited by M.C. Forde, London, UK, 11 pages [CD-ROM]. 176

PAGE 191

177 Boyd, A.J. (2000). Rehabilitation of reinforced concrete beams with sprayed glass fiber reinforced polymers. PhD thesis, Department of Civil Engineering, University of British Columbia, Vancouver, Canada. Collins, M. and Mitchell, D. (1991). Prestressed concrete structures, Prentice-Hall, Englewood Cliffs, New Jersey. Consolazio, G.R., Fung, J., and Ansley, M. (in press). M-phi-P diagrams for concrete sections under biaxial flexure and axial compression. American Concrete Institute (ACI) Structural Journal. International Federation for Structural Concrete (FIB) (2001). Technical Report on the design and use of externally bonded fibre reinforced polymer reinforcement (FRP EBR) for reinforced concrete structures. Bulletin 14, International Federation for Structural Concrete (FIB), Lausanne, Switzerland. Florida Department of Transportation (FDOT) (1998). Trucking manual, Office of Motor Carrier Compliance, Fifth edition, Tallahassee, Florida. Florida Department of Transportation (FDOT) (1999 revised 2002). Traffic engineering manual, Topic Number 750-000-005, Tallahassee, Florida. Fung, J. (2002). M-P-diagrams for prestressed concrete sections with surface bonded FRP composites. Masters thesis, Department of Civil and Coastal Engineering, University of Florida, Gainesville, Florida. Klaiber, F.W., Wipf, T.J., Russo, F.M., Paradis, R.R., and Mateega, R.E. (1999). Field/laboratory testing of damaged prestressed concrete girder bridges. Iowa Department of Transportation Report HR-397, Department of Civil and Construction Engineering, Iowa State University, Ames, Iowa. Labossieere, P., Neale, K.W., Rochette, P., Demers, M., Lamothe, P., Lapierre, P., and Desgagne, G. (2000). Fibre reinforced polymer strengthening of the Sainte-Emelie-de-lEnergie Bridge: design, instrumentation, and field testing. Canadian Journal of Civil Engineering, Vol. 27, 916-927. Mayo, R., Nanni, A., Watkins, S., Barker, M., and Boothby, T. (1999). Strengthening of bridge G-270 with externally bonded CFRP sheets. Research Investigation Number 98-012, Center for Infrastructure Engineering Studies, University of Missouri-Rolla, Rolla, Missouri. Meier, U. and Winistorfer, A. (1995). Retrofitting of structures through external bonding of CFRP sheets. Non-Metallic (FRP) Reinforcement for Concrete Structures, Proceedings of the Second International RILEM Symposium, E & FN Spon, London, UK, 509-516. Nawy, E.G. (2000). Prestressed concrete, Prentice Hall, Englewood Cliffs, New Jersey.

PAGE 192

178 Precast/Prestressed Concrete Institute (1999). PCI design handbook, Fifth Edition., Chicago, Illinois. Scheibel, S., Parretti, R., and Nanni, A. (2001). Repair and strengthening of impacted PC girders on bridge A4845. Missouri Department of Transportation (MoDOT) Report RDT01-017/RI01-016, Jefferson City, Missouri. Sen, R. and Liby, L. (1994). Repair of steel composite bridge sections using carbon fiber reinforced plastic laminates. Report Number 0510616, Department of Civil Engineering and Mechanics, University of South Florida, Tampa, Florida. Shahawy, M. and Beitelman (1999). Static and fatigue performance of RC beams strengthened with CFRP laminates. Journal of Structural Engineering, ASCE, 125(6), 613-621. Shanafelt, G.O., and Horn, W.B. (1980). Damage evaluation and repair methods for prestressed concrete bridge members. NCHRP Report 226, Transportation Research Board, National Research Council, Washington D.C. Shanafelt, G.O., and Horn, W.B. (1985). Guidelines for evaluation and repair of prestressed concrete bridge members. NCHRP Report 280, Transportation Research Board, National Research Council, Washington D.C. Spadea, G., Swamy, R.N., and Bencardino, F. (2001). Strength and ductility of RC beams repaired with bonded CFRP laminates. Journal of Bridge Engineering, ASCE, 6(5), 349-355. Tedesco, J.W., Stallings, J.M., and EL-Mihilmy, M. (1998). Rehabilitation of a reinforced concrete bridge using FRP laminates. Highway Research Center Report Number 930-341, Harbert Engineering Center, Auburn University, Auburn, Alabama. Teng, J.G., Smith, S.T., Lam, L., and Chen, J.F. (2001). Behaviour and strength of RC structures strengthened with FRP composites. Proceedings of the Ninth International Conference on Structural Faults and Repair, edited by M.C. Forde, London, UK, 15 pages [CD-ROM]. Teng, J.G., Chen, J.F., Smith, S.T., and Lam, L. (2002). FRP strengthened RC structures, John Wiley & Sons, LTD, Chicester, West Sussex, England. Tumialan, J.G., Huang, P-C, and Nanni, A. (2001). Strengthening of an impacted PC girder on bridge A10062, St. Louis County, Missouri. Research, Development and Technology Report Number 01-013, Center for Infrastructure Engineering Studies, University of Missouri-Rolla, Rolla, Missouri.

PAGE 193

179 Zobel, R.S., Jirsa, J.O., Fowler, D.W., and Carrasquillo, R.L. (1996, revised 1997). Evaluation and repair of impact-damaged prestressed concrete bridge girders. Center for Transportation Research Report 1370-3F, Bureau of Engineering Research, University of Texas at Austin, Austin, Texas.

PAGE 194

BIOGRAPHICAL SKETCH The author was born July 1, 1979, in Jacksonville, Florida to Bruce and Evelyn Lammert. She graduated from Samuel W. Wolfson High School in Jacksonville, Florida in 1997 and then moved to Gainesville, Florida, in the fall of 1997 to start her college career at the University of Florida. After receiving the degree of Bachelor of Science in Civil Engineering (with honors), she enrolled in the graduate program of the University of Florida Department of Civil Engineering to pursue a Master of Engineering with a focus in structural design. The author was initiated into Chi Epsilon, a Civil Engineering Honor Society, in December 2002. Her graduate career was focused on fiber reinforced polymer (FRP) repair of impacted bridge girders. She plans to receive her Master of Engineering degree in December 2003 after which she will pursue a career in structural design. 180


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Title: Structural evaluation of impact damaged prestressed concrete I girders repaired with Fiber Reinforced Polymer (FRP) materials
Physical Description: Mixed Material
Creator: Lammert, Kimberly Ann ( Author, Primary )
Publication Date: 2003
Copyright Date: 2003

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Title: Structural evaluation of impact damaged prestressed concrete I girders repaired with Fiber Reinforced Polymer (FRP) materials
Physical Description: Mixed Material
Creator: Lammert, Kimberly Ann ( Author, Primary )
Publication Date: 2003
Copyright Date: 2003

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STRUCTURAL EVALUATION OF IMPACT DAMAGED PRESTRESSED
CONCRETE I GIRDERS REPAIRED WITH FIBER REINFORCED POLYMER (FRP)
MATERIALS

















By

KIMBERLY ANN LAMMERT


A THESIS PRESENTED TO THE GRADUATE SCHOOL
OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT
OF THE REQUIREMENTS FOR THE DEGREE OF
MASTER OF ENGINEERING

UNIVERSITY OF FLORIDA


2003

































Copyright 2003

by

Kimberly Lammert















ACKNOWLEDGMENTS

Completion of this thesis and the research associated with it would not have been

successful without the help and guidance of a number of individuals. The author would

like to thank the chair and members of her committee Dr. Perry S. Green, Dr.

H.R.Hamilton, and Dr. Andrew J. Boyd for their guidance and assistance.

Others deserving of thanks due to contributions of resources and knowledge are Dr.

Thomas Sputo, Tony Michael, Jeff Brown, Johnny Fung, Jennifer Burtz, Vanessa Grillo,

Koren Britt, and Viswanath Urala.

The author would also like to thank the Florida Department of Transportation for

their financial assistance.

Lastly, the author would like to thank all of her close friends and family. Their

enduring support proved to be of the utmost value.
















TABLE OF CONTENTS
Page

A C K N O W L E D G M E N T S ................................................................................................. iii

LIST OF TABLES .................. .................. ................. ............ .............. .. vii

LIST OF FIGURES .............. ................................. ............. ........... viii

A B S T R A C T .......................................... ..................................................x iii

CHAPTER

1 IN TR OD U CTION ............................................... .. ......................... ..

O objective of C current Study ............................................................................ ...... 3
R elevance of C current Study ........................................................................... ...... 4

2 PREVIOU S RESEARCH ........................................................................9

Fiber Reinforced Polymer Properties and Behavior.........................................9
C condition A ssessm ents .................. ....... ...... .............. ...... .......... ............... 13
Repairs with Fiber Reinforced Polymer Systems.....................................................16

3 DESCRIPTION OF EXPERIMENTAL STUDY ............................................... 49

Description of Test Specimens in Damaged State .................................................49
Test Setup for Four Point Bending Tests .............. .............................................50
Repair D esigns and D esign Properties ............................................ ............... 51
R epair A application Procedures ........................................................ ............. 52

4 EXPERIMENTAL TEST RESULTS.................................... ......................... 70

Test 1 U ndam aged Control Results ........................................ ........................ 71
Test 2 D am aged Control R results ........................................ .......................... 73
Test 3 R JW atson R epair R esults.................................................................... .. ..... 75
Test 4 Air Logistics Repair Results ................................ ......................... ........ 78
Test 5 U F Sprayed R epair R esults.................................... ........................... ......... 81
Test 6 Edge Composites Repair Results ............................................ ...............83
Evaluation of Experimental Observations ............ .............................................83
Effect of Fiber Reinforced Polymers on Stiffness.............................................83
Effect of Fiber Reinforced Polymers on Capacity .................. ................85









Effect of Fiber Reinforced Polymers on Strains on Bottom Centerline of Test
Specim ens ........................................... .. .............. ............ 87
Other Effects Initial Imperfections and Bond ..............................................87
Summary of Test Specimen Failure M odes ................................ ............... 88

5 ANALYTICAL TEST RESULTS ................................................. ............... 110

C alculation of Effective Prestress................................................... ..................... ..110
Theoretical Cracking Capacity for Test Specimens in Undamaged and Damaged
State w without Repair..................... ..... ......................................... .......... 114
Theoretical Ultimate Capacity for Test Specimens in Undamaged and Damaged
State w without R epair.................................... ...... .. ........ ............ ............... ... 115
Theoretical Ultimate Capacities of Test Specimens 3, 4, 5, and 6 ...........................118
Support Conditions ............... ...... .......................... ..... ... ............. ............ 120
Theoretical Midspan Deflection at the Theoretical Ultimate Capacity of the
Girder in the Undamaged State.......................... ... .............................. 121
Comparison of Theoretical and Experimental Capacities and Deflections............122
Computer Program Implementation for all Test Specimens ..................................125

6 DESIGN RECOMMENDATIONS AND PRELIMINARY DESIGN
SPECIFICA TION S .................. ...................................... .. .......... .. 133

7 SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS FOR FUTURE
R E SE A R C H ................................................................ 138

S u m m a ry .......................................................................................1 3 9
Conclusions....................... ..... ......... .... ..... ............... 139
Future Testing Recommendations for Measurements to Assess Adequacy and
Performance of FRP Repair.... .. .............................................. .. ............... 140

APPENDIX

A R E P A IR D E SIG N S ......................................................................... ................... 142

R JW atson R epair D esign .............................................. ................................ 142
A ir L ogistics R epair D esign ............................................. ............................ 143
UF R epair D esign-A s D designed ........................................ .......................... 154
U F R epair D esign-A s B uilt ............................................... ............ ............... 158

B THEORETICAL EVALUATION TOOLS .......................................................163

Effective prestress............. ...... ...... ....... .... .......... ............163
Cracking Capacity for Test Specimens in Undamaged and Damaged State ..........165
Ultimate Capacity for Test Specimens in Undamaged and Damaged State............. 167
Ultimate Capacity for Repaired Test Specimens 3, 4, 5, and 6..............................170
Support Conditions ................... .................. ........................... ...............174
Deflection at Ultimate for Test Specimen 1 .................................. ............... 175











R E F E R E N C E S ...................................... ........................................................... .. 17 6

BIO GR APH ICA L SK ETCH ............................................................................. 180
















LIST OF TABLES


Table pge

1-1: Summary of Bridges in Florida Damaged due to Vehicular Impact.........................5

2-1: Qualitative Comparison of Carbon, Aramid and Glass Fibers..............................29

2-2: Comparison of Load, Deflection, Stiffness, and Crack Width................................29

2-3: Comparison of Midspan Deflection Before and After FRP Application ...................30

2-4: Comparison of Reinforcing Bar Stresses Before and After FRP Application ..........30

2-5: Comparison of Initial Stiffness and Strength for Various Retrofits.........................31

2-6: Comparison of Stiffness Before and After Retrofit.................................................31

2-7: Comparison of Ductility Ratios ............ .... ............................. ............... 31

3-1: Comparison of Design Repairs for Test Specimens 3, 4, 5, and 6...........................57

4-1: Experimental Capacities, Deflections, Strains, and Initial Flexural Stiffness of All
Test Specim ens at Cracking Load ...................................................................... 90

4-2: Experimental Capacities, Deflections, and Strains of All Test Specimens at
M maximum Load .......... .. ......... ....... ..................... ........ 90

5-1: Comparison of Experimental and Theoretical Capacities of Test Specimens 1, 2, 3,
4, 5,and 6 at Cracking ........... ................................... .. ...................... 128

5-2: Comparison of Experimental and Theoretical Capacities of Test Specimens 1, 2, 3,
4, 5 and 6 at M axim um Load ...........................................................................128

6-1:Environmental Reduction Factor for Various FRP Systems and Exposure
C conditions .......................................... ........................... 137
















LIST OF FIGURES


Figure pge

1-1: Number of Bridges in NBI versus Vertical Clearance of Selected 79000 Bridges......6

1-2: Number of Bridges in NBI versus Vertical Clearance of Selected 79000 Bridges by
Functional C classification ......................................... ................................. 6

1-3: Chaffee Road Bridge Overall Damage Looking East, Westbound Traffic................7

1-4: Close-up of Damage on the West Fascia Girder in the Vicinity of the Diaphragm.....7

1-5: Close-up of Damage on the East Fascia Girder Showing Severed Prestressing
S tra n d s ......................... ...... ........... .............................. ................ 8

2-1: Stress-Strain Diagrams for Mild Steel and Carbon, Glass, and Aramid Fiber
R enforced Polym ers .................. ........................... ................ ...... 32

2-2: FRP Placement for Increase in Shear Capacity .................................32

2-3: Flexural Failure Modes for FRP Strengthened Beams................................... 33

2-4: Shear Failure M odes for FRP Strengthened Beams.............................. ............... 34

2-5: Load versus Midspan Deflection for Specimen 1 with 54 ksi yield steel ..................34

2-6: Load versus Midspan Deflection for Specimen 2 with 54 ksi yield steel ..................35

2-7: Load versus Midspan Deflection for Specimen 3 with 45 ksi yield steel ..................35

2-8: Load versus Midspan Deflection for Specimen 4 with 54 ksi yield steel ..................36

2-9: Load versus Midspan Deflection for Specimen 5 with 45 ksi yield steel ..................36

2-10: Load versus Midspan Deflection for Specimen 6 with 45 ksi yield steel ...............37

2-11: M oment Deflection Relationship for Six Specimens.............................................38

2-12: Moment versus CFRP Strain for Repaired Specimens..............................39

2-13: Moment versus Reinforcing Bar Strains for Six Specimens.............................. 40









2-14: Moment versus Concrete Compressive Strain for Six Specimens...........................41

2-15: Overall Damage to 1-680 over CR L34, near Beebeetown, Iowa ..........................42

2-16: Moment Deflection Relationships for Beams 1 and 2...........................................42

2-17: Experimental and Theoretical Load Deflection Relationship for Beams 1 and 2 ...43

2-18: Load Deflection Relationship of Fully-Wrapped Beams ...................................43

2-19: Relationship Between Number of FRP Layers to Moment Capacity at Steel
Y field and U ltim ate ........................... .. ......................... .............. ..... .......... 44

2-20: Load Deflection Relationships for Fully-Wrapped and Partially-Wrapped
Sections with 3 Layers of FRP ....................... ................. 44

2-21: Theoretical and Experimental Midspan Deflection for the Analyzed Sections .......45

2-22: Load Deflection Relationship for Beams 1, 2, 3, and 4 .......................................46

2-23: Damage to Bridge A10062, St. Louis County, Missouri ...................................46

2-24: Overall Damage to Bridge A4845, Jackson County Missouri ..............................47

2-25: Close-up Damage to Bridge A4845, Jackson County Missouri............. ...............47

2-26: Load Midspan Deflection Relationship for the Control, Repaired with Fabric,
and Repaired with Spray Specim ens ............... ........ ............. ...... ............. 48

3-1: Typical Profile and Cross-Sections of Test Specimens in Undamaged State ............58

3-2: Typical Cross-Section Details for Test Specimens 2, 3, 4, 5, and 6 ........................59

3-3: Typical Damage for Test Specimens 2, 3, 4, 5, and 6 Showing Removed Concrete
and Tw o Severed Prestressing Strands................................. ....................... 59

3-4: Stress-Strain Curves for Removed Strands from Test Specimen 2.........................60

3-5: Load, Shear, and Moment Diagrams for All Four-Point Bending Tests....................60

3-6: Typical Instrumentation for Four-Point Bending Tests........................................61

3-7: Test Specim en 3 R epair D esign ........................................... .......................... 63

3-8: Test Specim en 4 R epair D esign ........................................... .......................... 64

3-9: Spray Equipment: Gun (G), Catalyst (C), and Resin (R)...........................................65









3-10: Detailed View of Spray Gun: Resin (R), Catalyst (C), Spray (S), and Chopper
(C h ) .................................................................................................. . 6 5

3-11: Spray Technique Procedure Showing How the Glass Fibers, Resin, and Catalyst
for Resin are Com bined in M id-A ir .................................... .................................. 66

3-12: Test Specim en 5 R epair D esign ........................................ .......................... 67

3-13: FRP Sampling Cores for Test Specimen 5 .................................... ............... 68

3-14: Test Specim en 6 Repair D esign ........................................ .......................... 69

4-1: Test Specimen 1 Load versus Midspan Deflection Curve ......................................91

4-2: Test Specimen 1 M measured Deflection Profile................................ ............... 91

4-3: Test Specimen 1 Strain Profile at North Load Point...............................................92

4-4: Test Specimen 1 Strain Profile at Midspan................... ........... ............... 92

4-5: Test Specimen 1 Strain Profile at South Load Point...............................................93

4-6: Test Specimen 1 Load versus Measured Strain Along Bottom Centerline of
B e a m .............................................................................. 9 3

4-7: Test Specimen 2 Load versus Midspan Deflection Curve ......................................94

4-8: Test Specimen 2 M measured Deflection Profile................................ ............... 94

4-9: Test Specimen 2 Strain Profile at North Load Point...............................................95

4-10: Test Specimen 2 Strain Profile at M idspan................... ...................... .......... 95

4-11: Test Specimen 2 Strain Profile at South Load Point.................... ... .............96

4-12: Test Specimen 2 Load versus Measured Strain Along Bottom Centerline of
B e a m .............................................................................. 9 6

4-13: Test Specimen 3 Load versus Time Showing Loading and Unloading Cycles .......97

4-14: Test Specimen 3 Load versus Midspan Deflection Curve ......................................97

4-15: Test Specimen 3 Measured Deflection Profile ....................................................98

4-16: Test Specimen 3 Strain Profile at North Load Point..............................................98

4-17: Test Specimen 3 Strain Profile at Midspan...........................................................99

4-18: Test Specimen 3 Strain Profile at South Load Point..............................................99









4-19: Test Specimen 3 Load versus Measured Strain Along Bottom Centerline of
B e a m ......................................................................... 1 0 0

4-20: Test Specimen 3 During Loading Showing Crack Propagation on South End of
Girder at the Bottom Layer of Prestressing ....................................... ................. 100

4-21: Test Specimen 3 at Failure Showing Concrete Cover Separation at South End....101

4-22: Test Specimen 4 Load versus Midspan Deflection Curve ...................................101

4-23: Test Specimen 4 M measured Deflection Profile................................................... 102

4-24: Test Specimen 4 Strain Profile at North Load Point .....................................102

4-25: Test Specimen 4 Strain Profile at Midspan .................................................103

4-26: Test Specimen 4 Strain Profile at South Load Point.............................................103

4-27: Test Specimen 4 Load versus Measured Strain Along Bottom Centerline of
B e a m ......................................................................... 1 0 4

4-28: Test Specimen 4 at Failure Showing Separation of FRP ......................................104

4-29: Test Specimen 4 at Failure Showing Shearing of FRP Stirrup ............................105

4-30: Test Specimen 5 Load versus Midspan Deflection Curve ...................................105

4-31: Test Specimen 5 M measured Deflection Profile................................................... 106

4-32: Test Specimen 5 Strain Profile at North Load Point................... ...............106

4-33: Test Specimen 5 Strain Profile at Midspan........ ............................ 107

4-34: Test Specimen 5 Strain Profile at South Load Point............. ............... 107

4-35: Test Specimen 5 Load versus Measured Strain Along Bottom Centerline of
B e a m ......................................................................... 1 0 8

4-36: Test Specimen 5 at Failure Showing Rupture of FRP at Midspan......................108

4-37: Test Specimen 5 at Failure Showing Rupture of FRP at Midspan on Tensile
Face of Girder. ............. ....... ............................ ............ ... ........ 109

5-1: Geometric Properties of the Cross-Section a) Girder without the Slab; b) Slab;
and c) G irder w ith the Slab .............................................................................. 130

5-2: Steel Areas and Location for a) an Undamaged Section; and b) a Damaged
Section ........................ .......... .................................. ........... 131









5-3: Normalized Moment versus Deflection for Test Specimens 1, 2, 3, 4, and 5..........132

5-4: Moment Curvature Comparison of Theoretical Data from Computer Program and
E xperim ental D ata ............ ........................................................................ .... 132















Abstract of Thesis Presented to the Graduate School
of the University of Florida in Partial Fulfillment of the
Requirements for the Degree of Master of Engineering

STRUCTURAL EVALUATION OF IMPACT DAMAGED PRESTRESSED
CONCRETE I GIRDERS REPAIRED WITH FRP MATERIALS

By

Kimberly Ann Lammert

December 2003

Chair: Perry S. Green
Major Department: Civil and Coastal Engineering

In Florida, any vehicle over 13 feet 6 inches is defined as an over-height vehicle

and requires a permit to travel on any road system in Florida. Although regulations are

present and enforced that does not ensure that collisions with bridge structures due to

over-height vehicles will not occur. Impact damage due to over-height vehicles can lead

to structure collapse, reinforcement damage, girder misalignment, steel yielding,

connection failure, reinforcement exposure, concrete spelling, and concrete cracking.

After a collision has occurred in which an emergency repair is required, a fast,

inexpensive, effective, and easy repair is preferred.

The impetus for this research is to assist the Florida Department of Transportation

(FDOT) in establishing a quality products list (QPL) of acceptable methods for the repair

of impact damaged prestressed concrete bridge girders using fiber reinforced polymer

(FRP) materials. The necessity of a quality products list was realized after the Chaffee

Road Bridge over Interstate 10 in Jacksonville was severely damaged on July 6, 2001,









from an over-height vehicle and needed immediate repair. At that time, the FDOT had

only one company or method to repair the bridge, and because most FRP methods are

proprietary to a company, realized the necessity of needing other approved alternatives.

The existence of a quality products list will enable qualified companies to be certified by

the FDOT to perform emergency repairs to vehicular damaged bridge girders using FRP

materials.

Full-scale girder tests of six type II AASHTO girders were performed. The six

tests represented an undamaged control specimen, a control specimen with simulated

damage, and the four specimens with simulated damage that were then repaired with

different FRP systems. The FRP systems varied in material types, laminate properties,

application procedures, reinforcing schemes, and repair design procedures and

assumptions. The specimens were tested to failure to determine moment and shear

capacities, as well as deformation and ductility behavior of the undamaged, damaged, and

repaired girders. Through experimental and analytical comparisons of the behavior of the

repaired girders to the undamaged girder, the specific FRP systems were evaluated for

acceptance to the FDOT Quality Products List. The FRP systems were evaluated

structurally based on their ability to restore the undamaged shear and moment capacity of

the prestressed concrete girder, type of failure mode, cost of repair system including

installation, and ease of installation.














CHAPTER 1
INTRODUCTION

The 1990 edition of the American Association of State Highway and

Transportation Officials' A Policy on Geometric Design ofHighways and Streets states

that the minimum vertical clearance for highways is recommended as 14.5 feet and is

desired as 16.5 feet (AASHTO 1990). Not all overpass bridges currently meet these

standards either due to being constructed before these standards were issued or paving

overlays of the underlying roadway (Bridge Engineering Software and Technology

[BEST] 2001). Figure 1-1, assembled by the National Bridge Inventory (NBI), shows the

clearance heights for 79000 bridges that have service both over and under the structure in

the U.S.. Figure 1-2 shows the clearance heights of these 79000 bridges in terms of their

functional classification either interstates and freeways or other arterials, collectors, and

locals.

In Florida, any vehicle over 13 feet 6 inches is defined as an over-height vehicle

and requires a permit to travel on any road system in the state (Florida Department of

Transportation [FDOT] 1998). By obtaining a permit the driver of an over-height vehicle

is provided knowledge of how to get a the desired destination on roads with sufficient

vertical clearance, and the FDOT is provided information about who, where, and when

the road system will have an over-height vehicle in case of bridge damage. Low

clearance postings are placed in advance of every bridge or structure with a minimum

vertical clearance of 14 feet 6 inches or less (FDOT 1999 revised 2002). In addition, low









clearance postings are placed on the structure for every bridge or structure with a

minimum clearance of 13 feet 6 inches or less.

Although regulations exist and are enforced does not ensure that collisions with

bridge structures due to over-height vehicles will not occur. Impact damage due to over-

height vehicles can lead to structure collapse, reinforcement damage, girder

misalignment, steel yielding, connection failure, reinforcement exposure, concrete

spelling, and concrete cracking. After a collision has occurred in which an emergency

repair is required, a fast, inexpensive, effective, and easy repair is preferred. Table 1-1

lists the bridges in Florida that have been damaged due to vehicular impacts over the past

two years. Table 1-1 provides for each of the impacted bridges the approximate date of

impact, district number, intersecting roads, and if any other impacts due to over-height

vehicles have been reported.

Many research projects have been conducted to determine what criteria should be

used for the assessment of damaged concrete structures (Shanafelt and Horn 1980,

Shanafelt and Horn 1985, Zobel et al. 1996, Arockiasamy and Barbosa 2000).

Establishment of these assessments would provide the necessary means to ensure that if

the structure is to be repaired that all potential problems are investigated, and a proper

and effective repair could be installed.

Numerous experimental studies of laboratory repairs and in field repairs with fiber

reinforced polymer (FRP) systems on concrete structures have been conducted (Sen and

Liby 1994, Arockiasamy 1995, Tedesco et al. 1998, Klaiber et al. 1999, Mayo et al. 1999,

Shahawy and Beitelman 1999, Labossieere et al. 2000, Tumialan et al. 2001, Scheibel et

al. 2001, Boyd and Banthia 2001, Spadea et al. 2001). All of the experimental research









has shown a promising future for fiber reinforced polymer systems for the repair of

damaged or deteriorated concrete or steel structures.

Objective of Current Study

The impetus for this research is to assist the Florida Department of Transportation

in establishing a quality products list (QPL) of acceptable methods for the repair of

impact damaged prestressed concrete bridge girders using fiber reinforced polymer (FRP)

materials. The necessity of a quality products list was realized after the Chaffee Road

Bridge over Interstate 10 in Jacksonville was severely damaged on July 6, 2001 from two

over-height vehicles and needed immediate repair. The Chaffee Road Bridge, built in

1960, is comprised of five Type III AASHTO girders with a 7" deck and an original

underclearance of 15'-6". Figure 1-3 shows the overall damage done to the Chaffee

Road Bridge's prestressed concrete girders including entrance and exit damage in the

background and foreground of the picture, respectively. Figures 1-4 and 1-5 show more

detailed views of the damage where the vehicles exited from under the bridge overpass

and where the vehicle made first contact with the bridge overpass and the resulting

damage. These figures show significant concrete cover loss to all girders, cracks

extending into the web for all girders, and severing of prestressing strands to the exterior

girders. At that time, the FDOT had only one company or method to repair the bridge,

and being that most FRP methods are proprietary to a company, realized the necessity of

needing other approved alternatives. The existence of a quality products list will enable

qualified companies to be certified by the FDOT to perform emergency repairs to

vehicular damaged bridge girders using FRP materials.









Relevance of Current Study

Full-scale girder tests of six type II AASHTO girders were performed. The six

tests represented an undamaged control specimen, a control specimen with simulated

damage, and four specimens with simulated damage that were then repaired with

different fiber reinforced polymer systems. The FRP systems varied in material type,

laminate properties, application procedures, reinforcing schemes, and repair design

procedures and assumptions. The specimens were tested to failure to determine moment

and shear capacities, as well as deformation and ductility behavior of the undamaged,

damaged, and repaired girders. Through experimental and analytical comparisons of the

behavior of the repaired girders to the undamaged girder, the specific fiber reinforced

polymer systems were evaluated for acceptance to be placed on the FDOT Quality

Products List. The FRP systems were evaluated structurally on their ability to restore the

undamaged shear and moment capacity of the prestressed concrete girder, the type of

failure mode that occurred, the cost of repair system including installation and ease of

installation.







5



Table 1-1: Summary of Bridges in Florida Damaged due to Vehicular Impact
Date District Facility Carries Intersects Multiple hits
1/15/2002 2 1-95 SR 206 YES
7/6/2001 2 1-10 CR 115C (Chaffee Road) YES
2/16/2001 4 1-95 Linton Blvd (SW 12th St) NO
4/24/2002 4 1-95 SR 708 (Blue Heron Blvd) YES
8/13/2001 7 SR 618 (Crosstown EXPY) 34th St YES
9/5/2002 8 SR 91 TPK SR 826 Palmetto EXPY NO
5/29/2001 8 SR 91 TPK SR 91 Jupiter Interchange NO















, I'


crnl



v I
I



\ /
i /
V'


N


12 13 14 15 16 17 18 19 2
Vertical Clearance (feet)

Figure 1-1: Number of Bridges in NBI versus Vertical Clearance of Selected 79000
Bridges (BEST 2001)




6000 ----- :7.-. 1,' -._. -


I/ i A -
54000 .. ,. .:- .-I.. .CI, Ii
5000 A




S/ \/ \
z


16
Vertical Clearance (feel)


Figure 1-2: Number of Bridges in NBI versus Vertical Clearance of Selected 79000
Bridges by Functional Classification (BEST 2001)


9000

8000

7000


S56000
c.
5000

o 4000

z 3000


I
I
I
I

/


3000


E
3
Z 2000



1000



0


14 15


18 19


S,"i,',, .














r II,---


iT
m-


FluIIe 1-4


I ^^ ,) .__


Figure 1-3: Chaffee Road Bridge Overall Damage Looking East, Westbound Traffic


- :~~ "~ l~l.."i
Ia-.


bp-


Figure 1-4: Close-up of Damage on the West Fascia Girder in the Vicinity of the
Diaphragm


" ~71~~;Lifr;;':'l"''::-~~~~ ~ ''":'i "'''''-'rilu
r..... r :~
ir*;;


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A


Figure 1-5: Close-up of Damage on the East Fascia Girder Showing Severed Prestressing
Strands


Y:~-~,c~c~cc~----~-t
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- -














CHAPTER 2
PREVIOUS RESEARCH

Fiber Reinforced Polymer Properties and Behavior

Previous research has shown a promising future for the use of fiber reinforced

polymer systems as an efficient means to strengthen or retrofit concrete structures in

order to resist increased design loads or repair damage (American Concrete Institute

[ACI] 2002). Fiber reinforced polymer systems are lightweight, easy to install,

noncorrossive, and inexpensive, when compared to prior repair techniques such as

bonded steel plates, external post-tensioning, or concrete and steel jacketing

(International Federation for Structural Concrete [FIB] 2001). When determining

whether to repair or replace a structure, it is crucial that if failure of the FRP system

occurs that it does not result in failure of the structure.

A fiber reinforced polymer system consists of fibers and a polymer matrix, namely

a stress-bearing component, and a stress-transferring component (FIB 2001). Fibers are

typically made of carbon, glass, or aramid and all exhibit an almost linear elastic behavior

to failure (ACI 2002). Figure 2-1 compares the stress-strain curves for carbon, glass, and

aramid fiber reinforced polymer systems to that of mild steel. The high tensile strength

of an FRP system provides increases in axial strength by fully wrapping the cross-section,

in shear capacity by placing the fibers perpendicular to potential shear crack development

as shown in Figure 2-2, and in flexural capacity by placing the fibers in the tension areas.

FRP systems can be used not only to increase flexural strength but can also be used to









increase the ductility in columns by providing confinement through wrapping the entire

cross-section of the column.

Polymer matrices or adhesives are available in epoxies, polyesters, and vinyl-esters.

Each type of fiber and adhesive has its advantages and disadvantages, which are

dependent on application procedure, type of structure, reason for rehabilitation, and the

configuration of the system. Table 2-1 provides a qualitative comparison of different

fiber composites. An effective fiber reinforced polymer system consists of fibers and an

adhesive working together so that the fibers can take on load from the original structure

through an adhesive bond. If the bond is not capable of transferring the required load, the

necessary capacity of the fibers can not be developed.

Fibers can be manufactured in a continuous or discontinuous form (FIB 2001).

Also, the fiber strands can be manufactured into unidirectional, bi-directional, or

multidirectional patterns or orientations forming a fabric which is capable of resisting and

distributing load in the manufactured directions (ACI 2002). A fabric is the arrangement

of fibers bound together in two or more directions.

Fiber reinforced polymer systems are divided into their different application

techniques and can be classified as a wet-lay-up system, pre-impregnated system, pre-

cured system or a sprayed system (ACI 2002, FIB 2001). A wet-lay-up system consists of

applying a saturating adhesive in-place to a dry reinforcement forming a cured laminate

on a structural member. A pre-impregnated system consists of an uncured fabric that has

been previously impregnated off-site with a saturating resin which is then bonded to the

surface of a structural member with or without an adhesive depending on system

requirements. A pre-cured system is applied using an adhesive to bond the off-site









previously cured and formed laminate to a structural member. The spray technique

combines the fiber, resin, and catalyst for the resin at the time of application and is

sprayed directly onto the structural member (Boyd and Banthia 2001).

The placement of the system on a structure and the orientation of the fibers

determines the type and amount of enhanced capacity which the rehabilitated structure

will be capable of resisting (ACI 2002, FIB 2001). To resist moment in a flexural

member, the fibers or fabric should be placed in the tension areas of the structural

member and oriented in the direction of the highest tensile stresses. To resist shear in a

flexural member, the fibers or fabric should be applied to the sides of the member and

oriented either perpendicular to latent or visible shear cracks or at 450 from the

longitudinal axis of the member as shown in Figure 2-2. When increasing the moment

capacity of a structural member it is necessary to check that sufficient shear capacity is

present to resist the corresponding increase in shear loads or additional shear

reinforcement must be provided.

There are many possible failure modes for beams flexurally strengthened with FRP

systems. If the ends of the FRP system are sufficiently anchored to the tension side of a

flexural member, brittle failure will occur when the beam reaches its ultimate flexural

capacity through tensile rupture of the FRP or concrete crushing as shown in Figures 2-3a

and 2-3b, respectively (Teng et al. 2002). Figure 2-3c shows that a brittle failure in shear

can occur if the flexural capacity exceeds the shear capacity of the strengthened member.

Brittle and premature failure can result from debonding due to high interfacial stresses at

or near the ends of an FRP system as shown in Figures 2-3d and 2-3e, respectively.

Brittle and premature failure can also result from debonding due to high interfacial









stresses from a flexural or flexural-shear crack formed away from the ends of a FRP

system as shown in Figures 2-3f and 2-3g, respectively. In addition, other failure modes

are possible such as failure of the adhesive, slip at the concrete substrate to adhesive

interface and slip at the adhesive to fabric interface or by a combination of any of the

modes described above. If the retrofitted member is a prestressed element, additional

failure modes can occur such as a prestressing strand steel failure.

Possible failure modes for shear strengthened beams are shear tension with FRP

rupture, and shear tension without FRP rupture where the FRP debonds as shown in

Figures 2-4a and 2-4b, respectively (Teng et al. 2001). Both failure modes for shear

strengthened members are brittle that can lead to an abrupt and catastrophic failure.

Equations, examples, and methods for determining FRP contribution have been

developed and included in the ACI (2002) code for reinforced concrete but not for

prestressed concrete. The only guidelines are that strain compatibility regarding the state

of strain in the prestressed member should be used to calculate FRP capacity increases

and that rupture of prestressing strands should be considered as a failure mode (ACI

2002). The absence of code standards means that design repairs are being performed

under increased engineering intuition, which could lead to incorrect assumptions.

Guidelines for detailing have been developed to avoid bond-related failures.

Suggested details incorporate the use of mechanical anchorages to assist in stress transfer,

tapering FRP lengths for multiple ply laminates, and providing sufficient overlap for

splicing FRP plies (ACI 2002).









Condition Assessments

Shanafelt and Horn (1980) reported on the findings of an extensive research and

compilation of statistics of all cooperating state departments of transportation. The

responding departments of transportation showed that of the 23,344 prestressed concrete

bridges in those states, an average of 201 bridges were damaged each year. The most

significant finding was that over 80 percent of the damage to prestressed concrete bridges

was due to over-height vehicles. The investigation produced the necessity for a standard

method to evaluate damaged bridges, classify the amount and types of damage, and how

to make decisions on the repairs. The types of repairs investigated in the report were

repairs with externally bonded reinforcing bars, external post-tensioning using strands or

bars, mild steel external sleeves, and internal strand splices. These repairs were

investigated through analytical calculations on beams classified as severely damaged.

Shanafelt and Horn (1985) further investigated the findings of the research by

performing repairs in the laboratory with external post-tensioning, mild steel external

sleeves, and internal strand splices. The purpose of the experiments was to establish

standards for evaluating damage to bridges, determining the amount of damage, and

designing and applying appropriate repair techniques. Although some of the repair

techniques investigated are no longer used solely for repairs, internal strand splicing is

still currently being investigated as part of a potential repair technique system.

Zobel, Jirsa, Fowler, and Carrasquillo (1996, 1997 revised) reported on field

inspections and laboratory experiments of an impact damaged prestressed concrete

girder. The damaged fascia girder was removed from the bridge and used to evaluate

non-destructive tests on impact damaged girders, and evaluate repairs with internal strand

splicing and concrete patching. Inspection of the girder showed cracking extending into









the web with undamaged exposed prestressing strands. The non-destructive techniques

used to evaluate the amount and type of damage to the girder were the rebound hammer,

impact echo, and spectral analysis of surface waves. In addition to the above

assessments, the girder was load tested without any repairs up to service load. The

results showed that the girder had a lower flexural stiffness and had larger deflections at

locations of significant damage and higher strains in strands with cross-sections that had

significant damage. The girder was repaired using several cast-in-place repair methods:

latex-modified vertical overhead mortars, single-component latex-modified mortars, and

fiber reinforced silica fume modified mortars over different damaged sections. Epoxy

was then injected to all remaining cracks and voids. After the repairs were completed,

the girder was reevaluated with the rebound hammer which showed a sound repair and

with the impact echo and spectral analysis of surface waves which showed that structural

integrity was generally restored. The beam was then load tested again up to its service

load. The results indicated that the girder had a higher flexural strength though it still had

larger deflections at locations with significant damage. The slope of the load versus

tendon strain curve showed significant improvement after the repair was completed.

The authors also investigated the effectiveness of internal strand splicing. Four

strands were intentionally severed at midspan to be spliced with four different splicing

assemblies. To determine the effectiveness of the splice, the girder was taken through a

series of load tests in which all four strands had been severed, all four strands spliced,

two splicing assemblies removed, and with all four splices removed. The load tests

demonstrated that the strain in the undamaged strands decreased when the damaged









strands were spliced suggesting a redistribution of strain to the spliced strands and the

capability of the splices to restore service load capacity.

Arockiasamy and Barbosa (2000) assessed techniques for the condition evaluation

and repair of concrete bridges. The research provides condition evaluation, selection of

repair types, repair design procedures, repair construction procedures, and case studies of

implemented repairs for bridge deck, substructure and superstructure elements. The

effective types of repairs that were considered in this study for prestressed girders were

external post-tensioning, internal splicing of strands, metal sleeve splices, and

replacement.

The Bridge Engineering Software and Technology (BEST) Center (2001) reported

on a study in which the purpose was to determine the extent of over-height vehicle

collisions and identify ways to prevent over-height vehicle collisions. A survey was sent

to every state requesting the following: average clearance height of bridges, maximum

vehicle height, fines for over-height vehicles, clearance postings, and collision data.

Only 29 states replied to the survey. Of the responding states only 12 provided statistics

on over-height collisions. Only 17 states stated that they keep records of over-height

vehicle collisions. The researchers determined from the submitted data that the number

of over-height vehicles increased by 1.3% per year from 1995 to 1999. In Maryland, 309

of the 1496 susceptible bridges or 20 percent have been damaged from over-height

vehicles. Of the 309 bridges that have been damaged 106 or 34% cross over an interstate

highway. To increase bridge clearances seven states or 24 percent indicated that they had

ground down the pavement that resulted from repeated resurfacing. The state of Georgia









responded that a program was ongoing to raise all existing interstate bridges to a

minimum clearance height of 16'-6".

Repairs with Fiber Reinforced Polymer Systems

Although a significant amount of research has been conducted on fiber reinforced

polymer (FRP) strengthening systems of reinforced concrete elements, little research has

been conducted on the use of FRP for the repair of prestressed concrete elements. This

section contains the last 10 years of research concerning repairs with FRP systems. At

present, standards for the design, application, and configuration of fiber reinforced

polymer strengthening systems on prestressed concrete structures have not been

developed.

Sen and Liby (1994) reported on the feasibility of using CFRP laminates to

strengthen steel composite beams through experimental and analytical procedures. The

composite beams were made of W8x24 steel sections with a 28" by 4 1/2" concrete slab.

The specimens were subjected to two loading cycles in four-point bending. The first

cycle consisted of loading the specimen to yield of the tension flange to simulate distress.

The second cycle consisted of loading the specimen to failure after the specimen was

repaired with CFRP laminates along the tension flange. Clamps were installed at the

ends of the laminates to resist peeling stresses. The clamps were designed based on the

stresses determined from a finite element analysis in the laminate and adhesive. The test

specimens consisted of two different steel yield strengths of 45 ksi and 54 ksi and were

repaired with two different laminate thicknesses of 2mm and 5mm. Specimens 1 and 2

were 54 ksi yield steel repaired with 2mm thick CFRP laminates, Specimen 3 was 45 ksi

yield steel repaired with 5 mm thick CFRP laminates, Specimen 4 was 54 ksi yield steel

repaired with 5 mm thick CFRP laminates, and Specimens 5 and 6 were 45 ksi yield steel









repaired with 2 mm thick CFRP laminates. After testing Specimens 1, 2, and 3,

additional bolts were used to further resist the stresses due to shear in the adhesive.

These additional three rows of bolts were spaced 16 1/2" apart at each end of the CFRP

laminate. Figures 2-5, 2-6, 2-7, 2-8, 2-9, and 2-10 show the load deflection relationships

for Specimens 1, 2, 3, 4, 5, and 6, respectively. The experimental results showed that the

CFRP laminates significantly increased the ultimate capacity of the steel composite

sections. A higher capacity with a more ductile failure mode could also be reached by

properly anchoring the CFRP laminates with the additional bolts as compared to the

unanchored repaired specimens.

Arockiasamy (1995) studied the flexural behavior of rectangular reinforced

concrete beams with a varying number of CFRP plates on the tension face through

experimental tests and analytical procedures. Six concrete specimens were tested to

failure, one was a control with no FRP, one was strengthened with one layer of FRP, two

were strengthened with two layers of FRP, and two were strengthened with three layers

of FRP. The beam dimensions were 8" wide by 1'-0" deep and were 8'-0" long with a

span to depth ratio of four. The beams were all tested to failure in two-point loading, and

were instrumented with seven strain gauges on the bottom centerline of the beam, four

strain gauges on one side of the beam, and five LVDT's equally spaced along the length

of the beam. Table 2-2 compares the ultimate capacity, deflection at midspan, initial

stiffness, and midspan crack width for all of the specimens, with the last number in the

beam name indicating the number of layers of CFRP. Figure 2-11 shows the moment

deflection relationship for all of the specimens. All of the beams with FRP systems failed

due to concrete crushing at smaller deflections than the control specimen. The bonding









of the CFRP plates to the tension face increased the flexural capacity, increased the

flexural stiffness, reduced crack widths, and reduced the CFRP, rebar, and concrete

compressive strains at midspan as the number of FRP plates increased. Moment versus

strain relationships for the CFRP, steel reinforcing bar, and concrete are shown in Figures

2-12, 2-13, and 2-14 respectively.

Tedesco, Stallings, and EL-Mihilmy (1998) investigated the effects of

rehabilitating a deteriorated reinforced concrete bridge in the field with external bonding

of FRP plates. The reinforced T-beam bridge was located on State Highway 110 near

Union Springs, Alabama. The simple span bridge had significant flexural and shear

cracking before repair. All of the reinforcing laminates used for repairs consisted of

unidirectional fibers oriented parallel to the longitudinal axis of the plate. Three of the

four beams were retrofitted with carbon fiber reinforced polymers on the tension face to

increase flexural capacity and with glass fiber reinforced polymers on the sides to

increase shear capacity. The remaining beam was only retrofitted with carbon fiber

reinforced polymers on the tension face. Load tests were performed before and after the

retrofit was applied to determine what affects the application of the FRP strengthening

system had on the bridge. Retrofitting the bridge resulted in smaller strains in the

longitudinal reinforcement and lower midspan deflections as shown in Tables 2-3 and 2-

4. The beam with only FRP on the tension face exhibited the least reduction in strain and

deflection indicating the significant effect of the GFRP on the overall structural behavior.

To verify the effects that FRP laminates have on strengthening the bridge, a finite

element analysis was performed. Results from the finite element analysis showed close









correlation to the in-field load tests for reinforcement strains and midspan deflections

with an average percentage difference of 5.7% and 5%, respectively.

Klaiber, Wipf, Russo, Paradis, and Mateega (1999) reported on the repair of an

impact damaged prestressed concrete bridge in Iowa in which field and laboratory tests

were conducted. The impacted bridge carries Interstate 680 over County Road L34

located near Beebeetown, Iowa. The bridge consists of two separate structures of the

same design and construction; one carries traffic eastbound and the other carries traffic

westbound. Each bridge consists of a concrete deck set on eleven beams with varying

vertical clearances. The westbound bridge was damaged while the eastbound bridge was

undamaged, thus providing a means to compare in-place load tests on the damaged and

undamaged structures. The three northernmost girders of the westbound bridge showed

significant damage at their midspans ranging from exposed and slack multiple

prestressing strands, a pre-existing severed strand, web cracking, and spelling of

concrete. Figure 2-15 shows some of the damage to the bridge. As part of the study, the

eastbound and westbound bridges were static load tested to determine whether they had a

different response to the applied load and load distribution. Although the two

northernmost damaged girders did not require replacement, they were subsequently

removed from the bridge for laboratory testing of fiber reinforced polymer systems and

the third northernmost girder was repaired insitu. Static load tests were again performed

on the westbound bridge after the removed girders were replaced. The field test results

showed that the behavior of the eastbound and westbound bridges were now basically the

same.









Beam 1 was tested to failure as the control specimen. Beam 1 had two severed

strands, which caused the beam to fail in combined shear and slab crushing at a moment

of 2067 kip-ft and corresponding deflection of 8.62". Beam 2 was damaged by severing

strands to be equivalent to Beam 1. Beam 2 was then repaired with three longitudinal

carbon fiber reinforced plates with additional fiber stirrups along the length to restore the

loss of capacity due to the severing of two strands and loaded to failure. Beam 2

ultimately failed due to debonding of the carbon fiber reinforcing system at a moment

and corresponding deflection of 2480 kip-ft and approximately 6.4". Figure 2-16

provides a comparison of the moment deflection relationships for Beams 1 and 2, the

series B1W is for Beam 1 and all of the other series are for Beam 2. By taking the

experimental strain in the composite at ultimate and multiplying by the area of FRP

material, the researchers determined that the force in the FRP was equivalent to three and

a half 250 ksi strands with a diameter of 0.5". Only two strands had been severed,

therefore the design goal of restoring the lost tensile capacity of the damaged strands was

reached. A 12% increase in ultimate capacity of the repaired girder (Beam 2) over the

control specimen (Beam 1) and a 10% increase in ultimate capacity over the theoretical

ultimate capacity of an undamaged member was reported.

Mayo, Nanni, Watkins, Barker, and Boothby (1999) reported on the strengthening

of a simple span reinforced concrete solid slab bridge in Iron County, Missouri. The

bridge, located on a mining truck operation route, Route 32, was strengthened in order to

remove weight restrictions. Through a comparison of design trucks in use at the time of

design to the present and including the current condition of the bridge such as corrosion

damage and cracking, it was determined that a 19% increase in flexural strength would be









required to remove the weight limit posting. To ensure the effectiveness of the FRP

strengthening system, two full-scale reinforced concrete beams were constructed and

tested to failure. The beams were constructed to simulate the existing bridge conditions

and geometry. Beam 1 was a control beam with no FRP strengthening, and Beam 2 was

strengthened with a one-ply carbon fiber reinforced polymer system to verify that a 20%

increase in capacity could be restored. Beam 1 failed due to concrete crushing, and Beam

2 failed due to FRP rupture. Figure 2-17 provides a comparison of the experimental and

theoretical load deflection relationships for Beams 1 (before strengthening) and 2 (after

strengthening). The laboratory test results showed that a 27% increase in flexural

strength with a finer crack pattern was achieved over the control beam test result. The

durability of the bonding of the FRP system was also investigated through static and

dynamic tests, confirming that peeling of the FRP should not be an issue. The bridge was

instrumented with deflection gauges and then load tested while before and after FRP

strengthening. The in-field load tests indicated a slight increase in stiffness due to

decreased deflection of the rehabilitated bridge.

Shahawy and Beitelman (1999) studied the static performance of reinforced

concrete beams strengthened with CFRP sheets. Eight T-beams of approximately 19'-0"

long were tested to failure in four-point bending. One specimen was tested as a control

with no FRP reinforcement. Five specimens had FRP applied to the entire stem, meaning

the bottom and sides of the beam, with varying layers of FRP. For the five fully-wrapped

specimens, there was one specimen with one layer of FRP, one specimen with three

layers of FRP, one specimen with four layers of FRP, and two specimens with two layers

of FRP. The remaining two specimens had two layers FRP applied only to the bottom









surface of the stem. All of the specimens were repaired with unidirectional carbon fiber

sheets. The T-beams were instrumented with LVDT's at the load points, supports, and

midspan and with strain gauges at the load points and midspan. The moment versus

midspan deflection curve for the fully wrapped beams and control beam show an increase

in capacity and a reduction in deflection of all of the repaired beams when compared to

the control specimen shown in Figure 2-18. Figure 2-19 shows how the number of layers

of FRP affects the flexural capacity at steel yield and ultimate for the fully wrapped

beams. The partially wrapped specimens showed lower moment capacity and deflection

than the fully wrapped sections presented in Figure 2-20. The partially wrapped beams

failed prematurely due to concrete cover separation, demonstrating the significance of

providing sufficient anchorage through wrapping the beams. A two-dimensional finite

element program, which takes into account the non-linear response of the component

materials was implemented. Four cross-sections were analyzed with this program

representing the control specimen, the partially wrapped beam with two layers of FRP,

and the remaining were fully wrapped cross-sections with two and three layers of FRP.

The program resulted in very consistent values for midspan deflection for all cross-

sections with slightly higher deflections and capacities at ultimate as shown in Figure 2-

21.

Labossieere, Neale, Rochette, Demers, Lamothe, Lapierre, and Desgagne (2000)

reported on the strengthening of the Sainte-Emelie Bridge that needed upgrading in order

to withstand heavy loads due to a nearby timber distributor. The Sainte-Emelie Bridge

carries Route 131 over the Riviere Noire near Sainte-Emelie-de-l-Energie in Quebec,

Canada. The bridge is a one-span monolithic deck and T-section reinforced concrete









structure with no significant signs of deterioration such as concrete spelling or corrosion.

The weight of standardized design trucks has increased since the design of this bridge.

Therefore, to bring the bridge up to current standard design loads, the bridge required a

35% increase in moment and 20% increase in shear capacity. To ensure the

strengthening procedure would be sufficient, numerous laboratory tests were performed

on one-third scale test specimens and analytical models were evaluated. The laboratory

tests included different potential climatic effects, behaviors of FRP materials, assessment

of an assortment of strengthening configurations for the T-section, and the bonding

capacity of the FRP system to concrete. Four beams were constructed and tested to

failure. Beam 1 was a control beam with no additional reinforcement. Beam 2 was

reinforced with six layers of carbon fiber for flexural strengthening on the full length with

glass fiber stirrups at varying spacing with an additional continuous strip of glass at the

top of the stirrups for added anchorage. Beam 3 was reinforced with one layer of carbon

fiber for flexural strengthening on the full length with the glass fiber stirrup spacing of

Beam 2 with an additional continuous strip of glass at the top of the stirrups for added

anchorage wider than Beam 2. Beam 4 was reinforced with six layers of carbon fiber for

flexural strengthening, three extending the full length and three apporximatley 3'-0"

shorter on both ends with glass fiber stirrups at varying spacing and an additional

continuous strip of glass at the top of the stirrups for added anchorage wider than Beam

2.

The control beam (Beam 1) failed in shear. Beam 2 failed in shear due to

debonding of the FRP stirrups but with a 40% load capacity increase over Beam 1. Beam

3 failed when the flexural strengthening failed in tension with a 60% load capacity









increase over Beam 1. Beam 4 failed when the flexural strengthening ruptured in tension

with a 50% load capacity increase over Beam 1. Figure 2-22 provides the load deflection

relationships for Beams 1, 2, 3, and 4.

A larger deflection at failure was observed for the beams with FRP than the control

beam consistent with the analytical models. The laboratory experiments showed the

effectiveness of the FRP material on strengthening reinforced concrete structures. The

final retrofit design for the Sainte-Emelie Bridge included three layers of carbon fiber

oriented longitudinally for the full length of the beams. In addition, glass fiber stirrups

with fibers oriented vertically anchored at the top with an additional glass fiber oriented

longitudinally were applied to the beams. To determine the in-field effectiveness of the

FRP system, the bridge was instrumented with strain gauges primarily at midspan and

located on the reinforcing steel inside the beams, on the FRP strips. A displacement

sensor was also placed under each beam at midspan. The bridge was load tested before

and after the FRP application to observe the bridge behavior. Decreases in the midspan

strains and deflections were observed after the FRP application as hypothesized in the

design and demonstrated in the laboratory.

Tumialan, Huang, and Nanni (2001) reported on an in-field fiber reinforced

polymer repair of an impact damaged bridge by an over-height vehicle in Missouri. The

damaged bridge, Bridge A10062, is located at the interchange of Interstates 44 and 270 in

St. Louis County, Missouri. After removal of the loose concrete, inspection of the

prestressed concrete bridge girders showed that the exterior girder had sustained the most

damage with two of the twenty prestressing strands severed. Figure 2-23 shows some of

the damage to the bridge. From an analytical procedure it was determined that a 190 k-ft









moment capacity would be needed from the FRP system to restore the girder to its

original strength. The repair called for a two-ply unidirectional carbon fiber reinforcing

polymer system with stirrups along the fiber length. No load tests, field measurements,

or laboratory simulations were conducted. After the repair was completed, it was

inspected and any potential void areas that were identified were epoxy injected to ensure

that a complete bond was provided.

Scheibel, Parretti, and Nanni (2001) investigated the effectiveness of fiber

reinforced polymer strengthening of eleven prestressed concrete girders from another

impact damaged bridge in Missouri. The damaged bridge, Bridge A4845, is located over

Route 291 on Route 24 in Jackson County, Missouri. An inspection of the eleven girders

showed only concrete spelling and the exposure of reinforcing steel and prestressing

tendons. Further inspection revealed that only three girders had exposed prestressing

tendons with a maximum of two tendons exposed in one girder. Figures 2-24 and 2-25

show overall and detailed views of the damaged bridge. For the design of the

strengthening system, the exposed tendons were assumed to be 50% effective. From an

analytical procedure it was determined that a 187 k-ft moment capacity would be needed

from the FRP system to restore the girder to its original strength. The repair called for a

two-ply unidirectional carbon fiber reinforcing polymer system with stirrups along the

fiber length. Long-term durability was addressed by applying the repair in an area away

from the damage to perform bond pull-off and torsion tests over five years. Initial

durability tests indicated a strong bond between the concrete surface and the FRP. No

load tests, field measurements, or laboratory simulations were conducted to show that the

designed repair was sufficient to restore the girder's original load capacity.









Boyd and Banthia (2001) reported on a new application method of fiber reinforced

polymer strengthening systems in which the fibers and matrix resin are simultaneously

sprayed onto the member surface. The spraying equipment provides a two-dimensional

random distribution of fibers with adjustable fiber lengths and has the ability to build up

the FRP strengthening system to any desired thickness. To examine the effectiveness of

this new method, three reinforced concrete channel beams that had been removed from a

badly deteriorated bridge were tested under third point bending to failure. The damage to

the beams consisted of cracking, loss of concrete cover, and reinforcement corrosion,

which varied, in each specimen making direct comparisons invalid. From the three

specimens, one was treated as a control with no repair while the other two were both

repaired with E-glass fibers: one repaired with the spray technique and the other repaired

with a continuous fiber system. Both applications were completed under laboratory

conditions meaning that the beams were repaired while upside down, which is not

possible in the field. The spray consisted of E-glass fibers embedded in a matrix

consisting of a polyester resin and a catalyst for the resin. The researchers also conducted

coupon tests and determined that the strength of the fiber reinforced polymer system

increased with increasing fiber length. Figure 2-26 shows the load deflection relationship

for the control specimen, the fabric repaired specimen, and the spray repaired specimen.

The experiments showed a greater increase in stiffness with the spray repair over the

fabric repair and a significant increase in strength for both repair fabric types. Also, for

material costs only, the spray repair was determined to be less expensive than the

continuous fiber repair.









Spadea, Swamy, and Bencardino (2001) reported on the laboratory testing of

reinforced concrete beams repaired with carbon fiber reinforced polymer laminates

reporting on respect to ductility, strength and failure modes. Eleven rectangular

reinforced concrete beams were constructed. The beams were divided into three series,

one control beam for each series with no external reinforcing referred to as Al, A2, and

A3. Within a series the internal reinforcing remained constant and the external

reinforcing was varied. Beams 1.1 and 3.1 were strengthened longitudinally with one

sheet of CFRP and no additional reinforcement. Each series contained a beam (1.2, 2.2,

and 3.2) that was strengthened longitudinally with one sheet of CFRP, a wide stirrup at

each support, and four small stirrups spaced along the length of the beam. Beams 1.3 and

3.3 were strengthened longitudinally with one sheet of CFRP, a wide stirrup at each

support, two small stirrups at midspan, and three small stirrups at each load point with a

plate above. Beam 2.3 was strengthened longitudinally with one sheet of CFRP, a wide

stirrup at each support, and ten small stirrups spaced along the length of the beam. The

stirrups or anchorages were made of steel and applied to the beams with the same

adhesive used for the CFRP. The steel plate anchorages were designed to either

counteract stresses at the end of the plates and control bond slip between the CFRP plate

and concrete, restrain movement of the plate, or provide lateral confinement for the

concrete in compression. The beams were tested in four-point bending to failure to

determine what increases in strength, ductility, and stiffness could be attained from the

different reinforcing schemes. The control beams with no external reinforcing failed as

expected due to concrete crushing. Beams 1.1 and 3.1 with only one sheet of CFRP and

no anchorages, failed suddenly due to debonding of the CFRP at a capacity 60.7% and









30.8% greater than their corresponding control specimen. The other beams, with

longitudinal and varying external stirrup reinforcing, all resisted a higher load than their

respective control specimen and failed in a more ductile manner than the beam with only

longitudinal reinforcing. The beams with anchorage reinforcing also reached higher

strains in the concrete and in the CFRP than the beams with only longitudinal reinforcing.

Structural ductility was quantified for deflection, curvature, and energy as defined by

shown in Equations 2-1, 2-2, and 2-3.

A,
PA A (2-1)







P=E (2-3)


Where Au = midspan deflection at ultimate load, Ay = midspan deflection at yield, u =

curvature at ultimate load, 4y = curvature at yield load, Etot = area under the curve at

ultimate load, and Etot = area under the curve at ultimate load. In Equation 3, Etot is taken

as the entire area under the load deflection curve and Ey is taken as the area under the

load deflection curve up to yield. Table 2-7 provides the defined ductility ratios for all

test specimens. Comparing the defined ductility ratios of the control beam to the other

beams shows a significant reduction for all of the repaired beams and shows a more

significant reduction in ductility of the beams repaired with only longitudinal reinforcing

over the beams with longitudinal and stirrup reinforcing.









Table 2-1: Qualitative Comparison of Carbon, Aramid and Glass Fibers (Meier and
Winistorfer 1995
Cri n Fiber Composite Sheets Made of:
Criterion
Carbon Fibers Aramid Fibers E-Glass Fibers

Tensile Strength Very Good Very Good Very Good
Compressive Strength Very Good Inadequate Good
Young's Modulus Very Good Good Adequate
Long-Term Behavior Very Good Good Adequate
Fatigue Behavior Excellent Good Adequate
Bulk Density Good Excellent Adequate
Alkaline Resistance Very Good Good Inadequate
Price Adequate Adequate Very Good

Table 2-2: Comparison of Load, Deflection, Stiffness, and Crack Width (Arockiasamy
1995)
Midspan Midspan
Ultimate Deflection Initial Crack Width
# of Load % at Ultimate % Stiffness % at Ultimate %
Beam plates (kip) diff (in) diff (kip-in/in) diff (in) diff
S5-STL 0 13.442 1.65 674.685 0.0147
S5-PRE1 1 14.973 11.39 1.34 18.79 813.803 20.62 0.00294 80
S5-PRE2 2 19.878 47.88 0.994 39.76 980.725 45.36 0.0092 37.41
S6-PRE3 2 22.014 63.77 0.962 41.7 1004.555 48.89 0.00554 62.31
S6-PRE4 3 20.675 53.81 0.643 61.03 1011.465 49.92 0.00474 67.76
S6-PRE5 3 26.126 94.36 0.881 46.61 1197.95 77.56 0.00456 68.98










Table 2-3: Comparison of Midspan Deflection Before and After FRP Application
(Tedesco et al. 1998)
Before FRP After FRP Percent
Girder (mm) (mm) Difference
Loading Position 1
1 6 5.6 7
2 7.9 7.3 8
3 7 6.3 10
4 3.2 2.8 12
Loading Position 2
1 5.4 5.2 5
2 7.7 7.1 8
3 7.3 6.6 10
4 3.7 3.3 12
Loading Position 3
1 3.1 3 2
2 6.5 6 8
3 8.6 7.7 10
4 6.3 5.5 12
Loading Position 4
1 3.6 3.5 4
2 6.9 6.3 9
3 8.3 7.4 11
4 5.6 5 12

Table 2-4: Comparison of Reinforcing Bar Stresses Before and After FRP Application
(Tedesco et al. 1998)
Before FRP After FRP Percent
Girder (MPa) (MPa) Difference
Loading Position 1
1 83 77 7
2 91 85 7
3 82 74 10
4 37 34 9
Loading Position 2
1 75 72 4
2 88 82 7
3 84 77 8
4 45 41 9
Loading Position 3
1 39 37 4
2 72 66 8
3 106 94 11
4 82 74 10
Loading Position 4
1 47 44 6
2 76 69 9
3 102 90 12
4 73 66 10










Table 2-5: Comparison of Initial Stiffness and Strength for Various Retrofits (Boyd and
Banthia 2001)
Stiffness Strength
Retrofit Initial Change Peak Change
Type (kN/mm) (%) (kN) (%)
None 6.69 214
Fabric 7.67 15 284 33
Spray 9 35 419 96



Table 2-6: Comparison of Stiffness Before and After Retrofit (Boyd and Banthia 2001)


Stiffness
Retrofit Before Retrofit
Type (kN/mm)


Stiffness
After Retrofit
(kN/mm)


Fabric 6.9 7.67 11
Spray 7.78 9 16



Table 2-7: Comparison of Ductility Ratios (Spadea et al. 2001)


Al 54 6.2 8 11.6
A1.1 86.8 1.5 1.5 1.9
A1.2 98 4.3 2.9 9.5
A1.3 96.7 3.5 4 7.6
A2 29.2 7.5 11.5 17.4
A2.2 75 4.8 3 13.2
A2.3 77.7 5.8 4 16.1
A3 57.2 7 10.3 15.7
A3.1 74.8 2 2.2 3.2
A3.2 98.8 4.5 3.3 10.9
A3.3 98.3 3.8 4.9 8.4


Change
(%)


Failure Deflection Curvature
Load (kN) Ductility Ductility


Beam


Energy
Ductility











6



4 FR2



2


M' ild steel


0.02 0.04 S

Figure 2-1: Stress-Strain Diagrams for Mild Steel and Carbon, Glass, and Aramid Fiber
Reinforced Polymers (FIB 2001)


Fiber


Figure 2-2: FRP Placement for Increase in Shear Capacity










(a) FRP rupture





(b) \\
Concrete crushing /






Shear crack



(d) High-stress zone


CracK propagation


(e) Hgh-lires zone


S Crack propagation

(f) Load
Crack
High-stress zone N
Crack propagation



(g) Load


Figure 2-3: Flexural Failure Modes for FRP Strengthened Beams (Teng et al. 2002)












',, -', "' :i :', ,' I"1 ; "',,'ie r c' ,' \,
,", ,:P ,u tr st11," .. i h
,t ,, I ". .
. i. . ._. ., r crick.', ,"
II
*111
FRPrupture sat here
in i li ;: ti 1 % iu II',


'IFRP rapture starts here


Debonded zone

1 .I I ,:


, ,1 I L' f CI I I


Figure 2-4: Shear Failure Modes for FRP Strengthened Beams (Teng et al. 2001)


1 .0 2.0 3.0


DISPLACEMENT (in)

Figure 2-5: Load versus Midspan Deflection for Specimen 1 with 54 ksi yield steel (Sen
and Liby 1994)


0.0
0.0


































1.0 2.0 3.0


DISPLACEMENT (in)

Figure 2-6: Load versus Midspan Deflection for Specimen 2 with 54 ksi yield steel (Sen
and Liby 1994)


40



30



20


0
0.


0


DISPLACEMENT (in)

Figure 2-7: Load versus Midspan Deflection for Specimen 3 with 45 ksi yield steel (Sen
and Liby 1994)


.. .. .. . .. . . .. .. . .. . .
-


o:o:
.0.






-I


-.. Unreinforc ed
5mm CFRP Plate


10 -


0
0.0







36




80



60
5 0 ... ... -- .. ... .... ... .. .... .... ... ........ .... ..




S. -. .. -. ... -. . .
40

30 -

20 -
Unreinforced
0 .. 5mm CFRP Plate

0
0.0 1 .0 2.0 3.0 4.0
DISPLACEMENT (in)

Figure 2-8: Load versus Midspan Deflection for Specimen 4 with 54 ksi yield steel (Sen
and Liby 1994)



60





40 -................- ---


3 0 -. ..... . .. ............ .. .... ..
2 0 .... . . ... . . .. . .




20

Unreinforc ed
2mm CFRP Plate
0
0.0 1.0 2.0 3.0 4.0 5.0
DISPLACEMENT (in)

Figure 2-9: Load versus Midspan Deflection for Specimen 5 with 45 ksi yield steel (Sen
and Liby 1994)












70












20
U nreinf orc ed
5mm CFRP Plate
10
0 ---------------------






0.0 1 .0 2.0 30 4.0
DISPLACEMENT (in)

Figure 2-10: Load versus Midspan Deflection for Specimen 6 with 45 ksi yield steel (Sen
and Liby 1994)












600




500




S400

0.


S300
E
0
E


< 200




100


.. ..-I I I I I I I
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8
Midspan deflection (inches)


Figure 2-11: Moment Deflection Relationship for Six Specimens (Arockiasamy 1995)






39




600-




500-




,400




S300-




200- -- PRE1. One plate

S-c-0 PRE2, Two plates

100- -A- PRE3, Two plates

PRE4, Three plates

S-a- PRE5, Three plates

-2000 0 2000 4000 6000 8000 10000 12000
CFRP plate strain (ue)

Figure 2-12: Moment versus CFRP Strain for Repaired Specimens (Arockiasamy 1995)






40




600




500




,400-




300-
E
0


SC__ control beam
2 200-
1-+-- PREl, One plate

-- PRE2, Two plates
100- -- PRE3, Two plates

-- PRE4, Three plates

---- PRE5, Three plates

-2000 0 2000 4000 6000 8000 10000 12000 14000 16000
Steel strain (ue)

Figure 2-13: Moment versus Reinforcing Bar Strains for Six Specimens (Arockiasamy
1995)






















-40




0 3
E


-a
< 20




10


10



Control beam
)0
-- PRE1, One plate

--- PRE2, Two plates

0- -- PRE3. Two plates -

-a- PRE4, Three plates

-e- PRE5, Three plates
0


-14000 -12000 -10000 -8000 -6000 -4000 -2000 0
Concrete compressive strain (ue)

Figure 2-14: Moment versus Concrete Compressive Strain for Six Specimens
(Arockiasamy 1995)


2000








42
































Figure 2-15: Overall Damage to 1-680 over CR L34, near Beebeetown, Iowa (Klaiber et
al. 1999)



2,500 -


2,000 -x x x -x
xx
& |? | x"' k "


S1,000 -
L I /x
,. x x \


x *--West 1 14 Pt
x
S K ----Midspan

x- B1W
xx
-------------- -n ----------- ------------ ---West --4--t



0 2 4 6 8 10 12
Deflection, in.


Figure 2-16: Moment Deflection Relationships for Beams 1 and 2 (Klaiber et al. 1999)





































0 1 2 3 4 5 6 7 8

Deflection (in.)
Figure 2-17: Experimental and Theoretical Load Deflection Relationship for Beams 1
and 2 (Mayo et al. 1999)


Midipan Deflection (mm)
40 60


2000 ------ -- -
| 21500--1* .....

' 15jO
a



0 "





0 0.5 1 13 2 7.5
Midspan Deflk(cn itcnris


200




I Layer

2 Layers

Layers 50

S 4 Layrn

3 3-5 4 4.5


Figure 2-18: Load Deflection Relationship of Fully-Wrapped Beams (Shahawy and
Beitelman 1999)


















I1.s -- ... .-------..- ---.-- .

a 1.7 At Ultimate















Number of Layers

Figure 2-19: Relationship Between Number of FRP Layers to Moment Capacity at Steel
Yield and Ultimate (Shahawy and Beitelman 1999)
1.5-

l2


Dc~nlctwl (mm)
Mi


S 0,5 1 1.5 2 2.5
DOfrllD D (1ncbl )


00




-- W2 4-A
1 '- 0


3 3,5 4 4.5


Figure 2-20: Load Deflection Relationships for Fully-Wrapped and Partially-Wrapped
Sections with 3 Layers ofFRP (Shahawy and Beitelman 1999)


300


2500




15000


5o Ioo
1000













MlMdspn Deflectton (mmn)
0 20 40 60 80 100
p- -


2500 I


S
6



S1500

so n
.- 1000


500


0






3000


2500


2000M

S1500
1 000






0


0 05 1 1.5 2 2,5 3 33 4
Mldpsan Dellecllon inchess)

Midspan Deflection (mm)
0 20 40 60 80 100


C-OLS



- Computer Model "







I I


0 0.5 1 15 2 2, 3 3,5 4 4,5 0.5 1 15 2 25 3 3.5 4 4
Midspan Deflection (inches) Midspan Deflection (inches)


Midspan Deflection (mm)
0 20 40 60 80 100


P-2L


-





Con pter Mobdel
4 -














1 ... mi er IM
0.5 1 .5 2 2.5 3 3.5 4 4
Midspen Deflection achese)

Ntdspsa Deflectio (ram)
0 20 40 64 80 100


Figure 2-21: Theoretical and Experimental Midspan Deflection for the Analyzed Sections
(Shahawy and Beitelman 1999)


300




200
C


S
ISO











100
so
0
50
S
5







300

250 ?

200 -
E
150
*1


50

0
5


t








350

300

S250

S200
a.
0
loo


0 25 50 75 100 125 150 175
Midspan deflection (mm)


Figure 2-22: Load Deflection Relationship for Beams 1, 2, 3, and 4
2000)





mm o.


(Labossieere et al.


Figure 2-23: Damage to Bridge A10062, St. Louis County, Missouri (Tumialan et al.
2001)

































Figure 2-24: Overall Damage to Bridge A4845, Jackson County Missouri (Scheibel et al.
2001)


Figure 2-25: Close-up Damage to Bridge A4845, Jackson County Missouri (Scheibel et
al. 2001)
al. 2001)


... .. *', '": : ,. '"" """...' ....... .. ',^": .



.."B"8. **_ -
: .. ... .. '













400

SSpray

300



200 rtro /l



100




0 25 50 75 100 125 150
Def)ction (mn)
Figure 2-26: Load Midspan Deflection Relationship for the Control, Repaired with
Fabric, and Repaired with Spray Specimens (Boyd and Banthia 2001)














CHAPTER 3
DESCRIPTION OF EXPERIMENTAL STUDY

In order to test different repair methods several identical test specimens were

needed that were in similar conditions. Six unutilized prestressed concrete type II

AASHTO bridge girders were found in a pre-cast yard in south Florida. The 43'-9"

girders were thought to be approximately 26 years old and to have the same constituents

based on corresponding lengths, strand patterns, and stirrup locations. These six girders

provided the testing of four repair methods leaving two girders as control specimens.

Two control specimens were needed to establish the moment and shear capacities at an

undamaged and damaged condition and therefore verify that sufficient additional moment

and shear strength would be attained from the various repair techniques to restore the

original moment and shear capacities of the girder.

Before any testing was initiated, a 1'-0" deep by 2'-0" wide slab having an average

6300 psi compressive strength was cast on each of the girders. The slabs were cast in

order to counteract the effect of camber due to prestressing and to duplicate in-field

conditions of having a deck or dead load acting on the girders. Figure 3-1 shows a

detailed cross-section and profile of the girders in the undamaged state.

Description of Test Specimens in Damaged State

All of the test specimens except for the undamaged control specimen were

damaged to replicate the most probable conditions that impact of an over-height vehicle

would cause. To simulate the impact damage in the laboratory, first a 5'-0" long section

of concrete was removed from each side of the bottom flange centered about midspan.









The hatched area in Figure 3-2 represents the removed concrete. After removal of the

concrete, four prestressing strands were exposed, two on each side, that were severed

representing an 18% loss in capacity. A concrete patch was then applied to the removed

concrete area to restore the original gross cross-section. An average 4500 psi concrete

compressive strength for all of the girders was determined by testing cores taken from

one test specimen. The location of the prestressing hold downs at the third points was

determined when the concrete was removed to damage the specimens. Direct tension

tests were performed of the removed 7/16" diameter, 250 ksi strands. Figure 3-2 shows

the cross-section of the girders at their midspans and ends and gives details as to the

damage executed on the girders. Figure 3-3 shows a typical girder in the damaged state

before the concrete patch was applied. The results of the direct tensile tests for the

removed strands are shown in Figure 3-4. The surfaces of all of the girders were

sandblasted on the sides and bottom in order that a good bond surface for the FRP repairs

would be produced.

Test Setup for Four Point Bending Tests

All girders were tested in a four point bending configuration where a uniform

moment exists between the load points and shear is only present between the load points

and reactions. Figure 3-5 shows the load diagram and resulting shear and moment

diagrams for the experiments. For the test setup, the load points were 10'-0" apart and

the supports were 40'-0" apart centered about the midspan. Figure 3-6 shows the test

setup for all test specimens and includes the locations of crack gauges, LVDTs, and load

cells. Figures 3-6b and 3-6c show the location of instruments corresponding to the

instrumentation lines shown in Figure 3-6a. The girders were loaded statically to failure

noting crack locations, sizes, and propagations during loading. The data from the tests









was reduced to determine the moment capacity, shear capacity, and midspan deflection of

the undamaged and damaged control specimens to be used for comparison and to later

verify if the FRP repair of the damaged girders had reached the required capacities. The

capacities would also provide the necessary values to be utilized in the design. The

supports for Test Specimen 4 were moved out an additional 6" on both ends due to the

termination of the FRP laminates occurring at the location of the supports for Tests 1, 2,

3, 5, and 6.

Repair Designs and Design Properties

The company that was to perform the repair was responsible for the design

calculations. No method or procedure, factors of safety, assumption for failure mode, or

fiber or resin material type was specified for the repair designs by the FDOT. Appendix

A contains the repair designs for Test Specimens 3, 4, and 5.

A comparison of the repair design properties and the corresponding description of

the repair for Test Specimens 3, 4, 5, and 6 is shown in Table 3-1. The design properties

in Table 3-1 for the repairs of Test Specimens 3, 4, and 6 are from data sheets provided

by the manufacturers through direct tensile coupon tests. ASTM D3039 (2000) gives

standards for determining tensile properties of polymer matrix composite materials. The

standards state that the extensometer gage length should be in the range of 0.5 to 2.0

inches. An extensometer gauge length of 0.5" was used for the direct tensile coupon test

data for Test Specimen 3. An extensometer gauge length of 2" was used for the direct

tensile coupon test data for Test Specimen 4. The design properties in Table 3-1 for the

repairs of Test Specimen 5 were taken from previous tensile coupon tests of

representative properties in which an extensometer gauge length of approximately 2









inches was utilized (Boyd 2000). For the Test Specimen 6 laminate properties, an

average of 0.12" elongation over an 8" long specimen was reported.

Repair Application Procedures

For Test 3, employees of RJWatson performed the installation procedure in which a

wet lay-up procedure was implemented. First a layer of epoxy was applied to the tension

face of the girder where the FRP strengthening system would be applied. After the epoxy

became tacky, one layer of pre-cut fabric was placed on the tension face of the girder and

rolled out to remove voids and prevent wrinkling of the fabric. Epoxy was then applied

to the exposed side of the fabric. When the exposed epoxy became tacky the next layer

of FRP was applied. This procedure was followed for each successive layer of FRP

applied. Figure 3-7 shows the installed repair for Test Specimen 3 in which the fabric

direction is oriented parallel to the longitudinal axis of the member. It should be noted

that all four layers are terminated at the same location.

For Test 4, the installation procedure was performed by Air Logistics in which pre-

cut fabric sizes of pre-impregnated resin were applied to the tension face with additional

stirrups of fabric placed to resist peeling of the fabric. A primer was first applied to the

tension side of the girder. When the primer became tacky, the fabric was unrolled

directly onto the surface. After two of the four layers were applied, water was sprayed

over the fabric with a garden hose activating the resin. The remaining two layers were

then applied and water was again sprayed over the fabric to activate the resin. As each

layer was applied, it was necessary to stretch the fabric in order to reduce air voids and

wrinkling of the fabric. After all of the layers were applied to the tension face, two layers

of bi-directional fabric were applied as stirrups following the above procedure with two

bolts placed through the ends into the web. Figure 3-8 shows the installed repair for Test









Specimen 4 in which the fabric direction is oriented parallel to the longitudinal axis of the

member and the layers are terminated at varying locations. In addition to longitudinal

reinforcement, a two-ply stirrup was installed just within the termination of the second

shortest layer to enhance the behavior of the FRP strengthening system. The stirrup

fabric was oriented at 0 and 900 as represented in Figure3-8 by the hatched area.

Magnum Venus Products (MVP) and graduate students from the University of

Florida performed the installation for Test 5. Figure 3-9 shows the equipment used for

the spray application procedure. The fibers in a roving form are routed into a chopper

mechanism attached to the spray gun. A detailed view of the gun is shown in Figure 3-

10. The chopping mechanism has two rollers with adjustable blades to vary fiber lengths

as the fiber roving passes through the rollers. The final FRP composite is comprised of a

resin, catalyst, and fibers that are fed separately into the spray gun and meet

simultaneously beyond the gun either in mid-air or on the application surface shown in

Figure 3-11. After an application of spraying was complete to an average thickness of 5

mils, ribbed metal rollers were used to compact the fibers, resin, and catalyst onto the

surface, remove air voids, and assist in bonding the material to the surface and underlying

layers. The above procedure was followed making each pass in which about 5 mils was

sprayed onto the surface and rolled out until the desired thickness was reached. The

repair was originally designed with fibers approximately 1.9 inches in length and its

representative properties with a laminate thickness of 0.35 inches. While attempting the

installation with the 1.9 inch fiber length, it was determined that it would be necessary to

reduce the fiber lengths. Since it was known from previous research that reducing fiber

length causes a reduction in composite strength, the repair was redesigned (Boyd 2000).









A new fiber length of 1.25 inches was utilized with a desired thickness of 0.50 inches to

produce an equivalent strength as specified in the original design. Figure 3-12 shows the

installed repair for Test Specimen 5 in which the multidirectional or randomly oriented

fibers were applied to the entire bottom of the girder and up 6" on both sides of the

bottom flange for 20'-0" centered about midspan. In addition a 2-0" wide stirrup of

randomly oriented fibers was applied up to the bottom of the top flange at the termination

of the repair at each end of the girder to enhance the shear strength of the repair and

ensure a good bond of the FRP strengthening system.

Difficulties were encountered during the two-day application of the spray FRP.

The spray application procedure had not been used in this type of application previously

where the spray had to be projected up onto the tension face of the beam simulating in-

field conditions with multiple layers having to be applied to reach the desired thickness.

In order to apply another layer, the underlying layer must be set such that the additional

layer does not pull the underlying layer off due to its self-weight. The FRP thickness

measurement device was incapable of piercing through the hardened underlying layers to

get accurate measurement of applied thickness so it was not known during or after the

application exactly what thickness was actually applied. Therefore, FRP cores were

taken from the specimen after loading. Figure 3-13 shows where samples were taken

from the cross-section and profile of the specimen.

A total of 73 samples were taken from the sides and bottom of the test specimen to

evaluate if the desired thickness was reached. The side samples are defined are those

taken from the web, and the angled and vertical side of the flange (numbers 1 through 5

and 9 through 13) as shown in Figure 3-13. The 73 samples yielded an average thickness









of 0.479" with a standard deviation of 0.196". Of the 21 samples taken from the tension

face of the girder none had a thickness equal to the desired thickness of 0.5". The

maximum and minimum thicknesses taken from the tension face of the girder were

0.368" and 0.139", respectively with an average thickness of 0.266" and a standard

deviation of 0.065". A total of 52 samples, 26 per side, were taken from the sides of the

test specimen with an average thickness of 0.565" with a standard deviation of 0.161".

There is a significant difference in thickness between the two sides of the specimen from

the two days of spraying in which only one side of the beam was repaired each day.

When the sides were evaluated by day, the results of the first day showed an average

thickness of 0.506" with a standard deviation of 0.102", and the results of the second day

had an average thickness of 0.625" with a standard deviation of 0.188".

The average thickness, as previously stated, for all of the 73 specimens was 0.479".

This was reasonably close to the desired thickness of 0.50", although it is apparent from

the bottom thickness measurements that the desired 0.50" was not reached during the

application of the FRP to the tension face of the girder.

The application procedure for Test Specimen 6 was performed by Edge Composite

employees, in which a wet lay-up procedure was implemented. All surface voids, in

which any FRP strengthening system was to be applied, were filled with a West System

Filler Epoxy containing micro-fibers to form a flat surface for a maximum bond. Then a

layer of resin was applied to the tension face of the beam in which the fabric would be

applied. When the resin became tacky the first layer of pre-cut fabric was applied to the

girder. A plastic trowel or scraper was used to remove any excess resin and to obtain the

desired constant thickness of resin between the layers. Each layer of fabric was rolled out









to remove voids and prevent wrinkling of the fabric. Resin was then applied to the

exposed side of the first layer of fabric. The procedure described above was followed for

the two successive layers. After the three layers were applied to the tension face, a two

ply stirrup was applied at the end of the longitudinal fabric to provide resistance against

peeling of the longitudinal fabric. The stirrups were applied to the girder using the above

described procedure. Figure 3-13 shows the installed repair for Test Specimen 6 in which

the fabric direction is oriented parallel to the longitudinal axis of the member and the

layers are terminated at the same location. In addition to the longitudinal reinforcement,

a two-ply stirrup was installed just within the termination of the fabric to enhance the

behavior of the FRP strengthening system. The stirrup fabric was unidirectional and

oriented perpendicular to the length of the girder as represented in Figure3-14 by the

striped area.










Table 3-1: Comparison of Design Repairs for Test Specimens 3, 4, 5, and 6
Test Number 3 4 5 6

Edge
Company RJWatson Air Logistics UF Edge
Composites
Carbon uni-
Carbon uni-
Fiber Material and directional with dl w h E-Glass multi- Carbon uni-
directional with
Description aramid cross directional fibers directional fibers
cross fibers
fibers

Resin Material Epoxy Polyurethane Polyester Epoxy

Elongation at 1.32%
1.21% 1.05% 1.50%
Failure 1.43%*
Tensile Modulus 1711
10500 11000 10100
(ksi) 1522*
Ultimate Tensile 15.66
127 115 150
Strength (ksi) 15.08*
Total Laminate 0 0.35
0.16 0.124 0.069
Thickness (in) 0.50*
Number of
4 4 1 3
Laminates
20 40 20
Total Length of 20 32
20 20
Laminates (ft) 20 24 20
20 16
Width of
16 12 18 18
Laminates (in)
2 plies of carbon 0.5 inches thick 2
2 plies of carbon
oriented at 0 and feet wide up to 2 plies of carbon
Stirrup Description N/A t uni-directional up
90, 12 ft from bottom of the
to bottom flange
midspan flange
* As-built properties
















2 '- -48 ~ _
S-- _Stirrup al
\ / Midspan
\// \\18 straight strands
Stirrups i :_'- _-_ 6'
6 draped strands with h
hold-down points at 1/3 span hold-down point


a) Typical Profile of Test Specimens in Undamaged State


24"
#5 bars at
mid-height of
slab ,= 60 ksi 1

#4 double bars
,= 60ksi


All strands
3" 7/16" 250 ksi
2" x 2" grid




Section at Midspan Section at End


b) Typical Cross-Sections of Test Specimens in Undamaged State

Figure 3-1: Typical Profile and Cross-Sections of Test Specimens in Undamaged State














#5 bars atmid-height
of slab f,= 60 ksi


#4 double bars f,= 60ksi


Removed portion of
concrete 5'-0" long -
centered about midsp an
(not in Test 1)


(not in Test 1)
Section at Midspan


All strands
7/16" 250ksi


Section at End


Figure 3-2: Typical Cross-Section Details for Test Specimens 2, 3, 4, 5, and 6


Figure 3-3: Typical Damage for Test Specimens 2, 3, 4, 5, and 6 Showing Removed
Concrete and Two Severed Prestressing Strands










300


250


200


, 150
Q)

100


50


0


-1% 0%


Strain
Figure 3-4: Stress-Strain Curves for Removed Strands from Test Specimen 2


P/2 P/2


P/2 15' 10' 15' P/2


P/2 P/2


P/2 15'- 6 10' 15'-6, P/2


For Test
Specimens
1,2,3,5, and 6


For Test
Specimen 4


pmm^


V
(kips)


M
(kip-ft)


Figure 3-5: Load, Shear, and Moment Diagrams for All Four-Point Bending Tests


/Chuck Failure

SChuck
Failure


Chuck
Failure


Strand 1
-- Strand 2
Strand 3
Strand 4



































NORTH


SOUTH


Figure 3-6: Typical Instrumentation for Four-Point Bending Tests a)Typical Instrumentation Layout for Test Specimens 1, 2, 3, 4, 5, 6


LEGEND


LVDT 2" Crack Load
Gauge Cell



























....... J J- --t ........ I J -----,

Instrumentation Line 'Instrumentation Line


LEGEND


LVDT __ 2" Crack Load
Gauge Cell



Figure 3-6(continued): Typical Instrumentation for Four-Point Bending Tests b)Typical Instrumentation Locations for Section A-A
Cross-Section in Figure 3-6a; c) Typical Instrumentation Locations for Section B-B Cross-Section in Figure 3-6a





















-Reaction Line


.1 0'


zj1 '-4"
Section at
Midspan


Figure 3-7: Test Specimen 3 Repair Design

























Reaction
Line-
A 12 Section at Section
-1 6' Midspan A-A
20'




Figure 3-8: Test Specimen 4 Repair Design


























Figure 3-9: Spray Equipment: Gun (G), Catalyst (C), and Resin (R)


S


ICh


Figure 3-10: Detailed View of Spray Gun: Resin (R), Catalyst (C), Spray (S), and
Chopper (Ch)




























Figure 3-11: Spray Technique Procedure Showing How the Glass Fibers, Resin, and
Catalyst for Resin are Combined in Mid-Air















-i









Section at Section
Midspan A-A


-4


Figure 3-12: Test Specimen 5 Repair Design














-1' 4- 5, '
,,L -


---I
I I I
i i I


Samples taken _
for numbers 1-13


SSamples taken /
for numbers 4-10


Figure 3-13: FRP Sampling Cores for Test Specimen 5


Support


1 13
2 12
43 11
51 19
678


1 .o~


























'-6"' -
Section Section
at Midspan A-A o
-o


Figure 3-14: Test Specimen 6 Repair Design














CHAPTER 4
EXPERIMENTAL TEST RESULTS

This chapter contains the experimental data for the six, four-point bending tests

that were conducted. Chapter 5 will serve as a comparison of the experimental results

shown in this chapter to calculated theoretical values.

Loading was terminated for each specimen once either load cell showed a reduced

reading therefore, no post ultimate data is shown for any of the tests. The total applied

load is the sum of the two load cell measurements shown in Figure 3-6a. The moment in

the constant moment region is calculated by taking one-half the total applied load and

multiplying it by 15'-0" for Test Specimens 1, 2, 3, 5, and 6 and 15'-6" for Test

Specimen 4. Strain and deflection profiles are plotted at 25, 50, 75, 90, and 100 percent

of the maximum load attained during each experiment. Strain gauges that either were

initially defective, or after a certain load level no longer functioned properly, were

omitted from the strain profiles. Large increases in the strain readings for a particular

gauge can be attributed to crack formation occurring inside of the two-inch gauge length

causing increased strain due to stress redistribution. The experimental neutral axis depths

were determined by assuming a best-fit linear strain distribution through the cross-section

for a specified load level using all properly functioning strain gauges. In order to

compare strain values at a certain load level and strain increases between load levels, the

strain gauge closest to the prestressing steel, 45 inches down from the top of the slab, was

chosen for comparison between all test specimens. In a case where the strain value was

unavailable at the steel level the value was determined by linear interpolation between the









compressive strain measured 6" down from the top of the slab and the extreme tensile

strain measured 48" down from the top of the slab. If the extreme tensile strain value at

48" down from the top of the slab was also unavailable, the strain at 48" down was set

equal to the bottom strain being measured closest to the unavailable data at the same load

level.

Test 1 Undamaged Control Results

The load versus midspan deflection for Test Specimen 1 is shown in Figure 4-1.

The experimental load capacities at cracking and ultimate are 131 kips and 268 kips with

corresponding midspan deflections of 0.50 and 5.95 inches, respectively. The load

deflection relationship, as shown in Figure 4-1, is linear up to the cracking load and

nonlinear between the cracking load and maximum load. Symmetrical deflections

occurred about the midspan of the beam as shown in Figure 4-2, except for a slightly

larger deflection at the south load point at the maximum load, Pmax. Figure 4-2 also

shows similar increases in deflection between 0 and 25 percent and between 25 and 50

percent of the maximum load, with larger increases in deflection between 50 and 75

percent and between 75 and 90 percent of the maximum load, and a substantial increase

in deflection between 90 and 100 percent of the maximum load.

The strain profiles in Figures 4-3, 4-4, and 4-5 all show an upward movement of

the neutral axis from the initial location of 20.7 inches down from the top of the slab into

the slab at Pmax. The final depths of the neutral axis at the north load point, midspan, and

south load point locations are 11.27", 9.05", and 5.73", respectively, down from the top

of the slab at the maximum load. The variation in neutral axis depths can be attributed to

either a slight imbalance of 4 kips at the maximum applied load with a higher load at the

south load point location, larger deflections, slipping of prestressing strands, yielding of









the strands, or a compression failure all on the south end of the girder. Figures 4-3, 4-4,

and 4-5 also show a significant increase in strain between 90 and 100 percent of the

maximum load that can be attributed to the formation of multiple cracks or the further

opening up of existing cracks that developed under the applied load. Figure 4-3 shows

only slight increases in strain when compared to the strain increases of Figures 4-4 and 4-

5. Figures 4-4 and 4-5 show similar strain profiles for the midspan and south load point,

which is reasonable since both locations are in the constant moment region. Strain

increases between the load levels at the prestressing steel layer at the north load point

shown in Figure 4-3 are 245 x 10-6 in/in, 1741 x 10-6 in/in, 1933 x 10-6 in/in, and 3508 x

10-6 in/in. Strain increases between the load levels at the prestressing steel layer at

midspan shown in Figure 4-4 are 456 x 10-6 in/in, 3612 x 10-6 in/in, 2615 x 10-6 in/in, and

7240 x 10-6 in/in. Strain increases between the load levels at the prestressing steel layer

at the south load point shown in Figure 4-5 are 339 x 10-6 in/in, 3303 x 10-6 in/in, 2224 x

10-6 in/in, and 7680 x 10-6 in/in. Comparing the strain increases at the prestressing steel

level shows significantly larger increases in strain at the south load point than at the north

load point, and strain increases at the south load point similar to the increases at midspan.

The larger strains at the south load point location can be attributed to either unbalanced

load cells with a higher load at the south load point location, larger deflections, slipping

of prestressing strands, yielding of the strands, or a compression failure all on the south

end of the girder. Figure 4-6 shows the load versus strain data for the strain gauges

shown in Figure 3-6a along the bottom centerline of the test specimen. The figure shows

similar strain readings for the four gauges up to the cracking load, and similar strains up

to failure for the two strain gauges at 1'-0" from midspan and for the two strain gauges









3'-0" from midspan. The maximum recorded bottom centerline strain at Pmax was

0.018447 in/in and occurred at 3'-0" north of midspan.

Concrete crushing was not the failure mode as expected, which allows for slipping

of the strands to be assessed as the failure mode. The slightly larger deflections and

larger strains recorded at the south load point location provide an assumption that the

slipping of strands occurred on the south end of the girder. The shear, moment, and

midspan deflection recorded at failure for the undamaged girder were 268 kips, 2012 kip-

ft, and 5.95 inches, respectively.

Test 2 Damaged Control Results

The load versus midspan deflection for Test Specimen 2 is shown in Figure 4-7.

The experimental load capacities at cracking and ultimate are 112 kips and 210 kips with

corresponding midspan deflections of 0.54 and 6.42 inches, respectively. The load

deflection relationship, as shown in Figure 4-7, is linear up to the cracking load and

nonlinear between the cracking load and maximum load. Symmetrical deflections

occurred about the midspan of the beam as shown in Figure 4-8, except for a slightly

larger deflection at the south load point at the maximum load, Pmax. Figure 4-8 also

shows similar increases in deflection between 0 and 25 percent and between 25 and 50

percent of the maximum load, with larger increases in deflection between 50 and 75

percent and between 75 and 90 percent of the maximum load, and a substantial increase

in deflection between 90 and 100 percent of the maximum load.

The strain profiles in Figures 4-9, 4-10, and 4-11 all show an upward movement of

the neutral axis from the initial location of 20.7 inches down from the top of the slab into

the slab at the maximum load. The final depths of the neutral axis at the north load point,

midspan, and south load point locations are 10.78", 8.16", and 5.33", respectively, down









from the top of the slab at the maximum load. The variation in neutral axis depths can be

attributed to either a slight imbalance of 1 kip at the maximum applied load with a higher

load at the south load point location, larger deflection, slipping of prestressing strands,

yielding of the strands, or a compression failure all on the south end of the girder.

Figures 4-9, 4-10, and 4-11 also show a significant increase in strain between 90 and 100

percent of the ultimate load possibly due to the formation of multiple cracks or the further

opening up of existing cracks. The strain profiles in Figures 4-9 and 4-11 for the north

and south load point locations show similar strain values for all load increments. A

comparison of the strains in Figures 4-9, 4-10, and 4-11 show that there were larger

increases at midspan than at the load point locations for all profiles up to 0.90Pmax. This

can be attributed to the damage imparted to the specimen only in the midspan area prior

to testing. Strain increases between the load levels at the prestressing steel layer at the

north load point shown in Figure 4-9 are 254 x 10-6 in/in, 1364 x 10-6 in/in, 1121 x 10-6

in/in, and 11759 x 10-6 in/in. Strain increases between the load levels at the prestressing

steel layer at midspan shown in Figure 4-10 are 536 x 10-6 in/in, 6392 x 10-6 in/in, 16971

x 10-6 in/in, and 1140 x 10-6 in/in. Strain increases between the load levels at the

prestressing steel layer at the south load point shown in Figure 4-11 are 277 x 10-6 in/in,

1243 x 10-6 in/in, 800 x 10-6 in/in, and 11832 x 10-6 in/in. Comparing the strain increases

at the prestressing steel level shows larger increases in strain at the midspan location than

at either of the load point locations except for the last load step where similar increases in

strain at the north and south load point locations were recorded. Larger increases at the

midspan location than at the load point locations are reasonable since the midspan is

within the constant or maximum moment region and the load points are not located in the









constant moment region. Figure 4-12 shows the load versus strain for the strain gauges

shown in Figure 3-6a along the bottom centerline of the test specimen. Similar strain

readings were recorded up to the cracking load for the four strain gauges. The two strain

gauges at 1'-0" from the midspan show similar strains up to failure and the two strain

gauges 3'-0" from midspan show similar strains up to failure. Figure 4-12 also shows

higher strains for the strain gauges 1'-0" from midspan at the maximum load than the

strain gauges 3'-0" from midspan, which were located one foot from the end of the

damaged area. The maximum recorded bottom centerline strain at Pmax was 0.023338

in/in and occurred at 1'-0" south of midspan.

Concrete crushing was not the failure mode as expected, which allows for slipping

of the strands to be assessed as the failure mode. The slightly larger deflections and

larger strains recorded on the south load point location provide an assumption that the

slipping of strands occurred on the south end of the girder. The shear, moment, and

midspan deflection recorded at failure for the damaged girder were 210 kips, 1576 kip-ft,

and 6.42 inches, respectively.

Test 3 RJWatson Repair Results

Two cycles of loading and unloading of Test Specimen 3 were performed before

the specimen was taken to failure as shown in Figure 4-13. The data provided in Figures

4-15 through 4-18 was taken from the last series of loading. The two initial cycles of

loading caused a slight softening if the load deflection curve, meaning that there was

some measurable reduction in the stiffness of the specimen at the beginning of the

loading cycle to failure. The load versus midspan deflection for Test Specimen 3 is

shown in Figure 4-14. The experimental load capacities at cracking and ultimate are 125

kips and 244 kips with corresponding midspan deflections of 0.60 and 3.15 inches,









respectively. The experimental cracking load for Test Specimen 3 was attained during

the second series of loading which means that for the third or final loading series, the

cracking load had previously been exceeded. The load deflection relationship, as shown

in Figure 4-14 is still linear up to the cracking load and nonlinear between the cracking

load and maximum load. Symmetrical deflections occurred about the midspan of the

beam as shown in Figure 4-15, except for a slightly larger deflection at the south load

point at the maximum load. Figure 4-15 also shows similar increments of deflection

between 0 and 25 percent and between 25 and 50 percent of the maximum load, with

increasingly larger increments of deflection between 50 and 75 percent, 75 and 90

percent, and 90 and 100 percent of the maximum load.

The strain profiles in Figures 4-16, 4-17, and 4-18 all show an upward movement

of the neutral axis from the initial location of 20.81 inches down from the top of the slab

into the slab at Pmax. The final depths of the neutral axis at the north load point and

midspan locations are 10.98" and 10.52", respectively, down from the top of the slab at

the maximum load. The neutral axis depth at the south load point location can not be

determined due to an insufficient number of data points. The determined final depths of

the neutral axis for the north load point and midspan locations are similar with a slightly

smaller neutral axis depth for the midspan location. This difference can be attributed to

either the midspan location being within the constant or maximum moment region and

the north load point location being located at the transition between the moment gradient

and constant moment regions or a slight imbalance of 7 kips at the maximum applied

load with a higher load at the south load point location. The strain profiles in Figures 4-9

and 4-11 for the north and south load point locations show similar strain patterns for all









load steps with slightly larger strains at the north load point. A comparison of the strains

in Figures 4-9, 4-10, and 4-11 show that there were larger increases at midspan than at

the load point locations for all profiles. This can be attributed to the damage imparted on

the specimen only in the midspan area prior to testing. Strain increases between the load

levels at the prestressing steel layer at the north load point shown in Figure 4-16 are 254

x 10-6 in/in, 1610 x 10-6 in/in, 1210 x 10-6 in/in, and 753 x 10-6 in/in. Strain increases

between the load levels at the prestressing steel layer at midspan shown in Figure 4-17

are 503 x 10-6 in/in, 1454 x 10-6 in/in, 1118 x 10-6 in/in, and 1174 x 10-6 in/in. Strain

increases between the load levels at the prestressing steel layer at the south load point

shown in Figure 4-18 are 209 x 10-6 in/in, 810 x 10-6 in/in, 650 x 10-6 in/in, and 2347 x

10-6 in/in. The largest increases for the strain gauges near the level of prestressing were

located at midspan for the first two load steps of 0.25 and 0.50Pmax, the north load point

for the next two load steps of 0.75 and 0.90Pmax, and the south load point for the last load

step up to Pmax. The pattern of strain increase shows fairly symmetric strains about

midspan up to 0.50Pmax equal to 122 kips which is approximately equal to the

experimental cracking load of 125.28 kips. Also, the largest strain increase occurred

between 0.90Pmax and 1.00Pmax at the south load point, which is at the location where a

portion of the girder completely separated from the rest of girder at failure. Figure 4-19

shows the load versus strain data for the strain gauges shown in Figure 3-6a along the

bottom centerline of the test specimen. Similar strain readings were recorded for the two

strain gauges on the north end of the beam up to failure. The figure shows slightly larger

strains for the strain gauges located 3'-0" from midspan than the strain gauges 1'-0" from









midspan throughout loading. The maximum recorded bottom centerline strain at Pmax

was 0.005083 in/in and occurred at 3'-0" north of midspan.

The failure mode of Test Specimen 3 was concrete cover separation as shown in

Figures 4-20 and 4-21. A crack formed at both ends of the FRP plate at the bottom layer

of prestressing and propagated towards midspan. The severe cracking at the south end of

the girder caused a portion of the bottom flange of the girder up to the bottom layer of

prestressing to fall to the ground with the adhered FRP. The shear, moment, and midspan

deflection recorded at failure for the repaired girder were 244 kips, 1829 kip-ft, and 3.15

inches, respectively. The ultimate shear and moment capacity corresponds to restoring

91% of the experimental ultimate capacity of the undamaged girder. The measured

deflection at Pmax for Test Specimen 3 corresponds to a 47% reduction in the deflection of

the undamaged girder (Test Specimen 1) at its maximum load.

Test 4 Air Logistics Repair Results

The load versus midspan deflection for Test Specimen 4 is shown in Figure 4-22.

The experimental load capacities at cracking and ultimate are 111 kips and 202 kips with

corresponding midspan deflections of 0.47 and 1.90 inches, respectively. The load

deflection relationship, as shown in Figure 4-22 is linear up to the cracking load and

nonlinear between the cracking load and maximum load. Symmetrical deflections

occurred about the midspan of the beam as shown in Figure 4-23, except for a slightly

larger deflection at the south load point at the maximum load. Figure 4-23 also shows

similar increases in deflection between 0 and 25 percent and between 25 and 50 percent

of the maximum load, with slightly larger increases in deflection between 50 and 75

percent, 75 and 90 percent, and 90 and 100 percent of the maximum load.









The strain profiles in Figures 4-24, 4-25, and 4-26 all show an upward movement

of the neutral axis from the initial location of 20.76 inches down from the top of the slab

into the slab at Pmax. The final depths of the neutral axis at the north load point, midspan,

and south load point locations are 9.08", 8.18", and 14.62", respectively, down from the

top of the slab at the maximum load. It is reasonable that the neutral axis depth at

midspan would be less than the neutral axis depths at the load point locations due to the

midspan location being within the constant or maximum moment region and the load

point locations being located at the transition between the moment gradient and constant

moment regions. The difference of the neutral axis depths at the load point locations can

be attributed to either a slight imbalance of 1 kip at the maximum applied load with a

higher load at the north load point location, larger deflections, slipping of prestressing

strands, yielding of the strands, compression failure, or failure of the FRP all on the north

end of the girder. The strain profiles in Figures 4-24 and 4-26 for the north and south

load points show similar strain patterns for all load increments with slightly larger strains

at the north load point. A comparison of the strains in Figures 4-24, 4-25, and 4-26 show

that there were larger increases at midspan than at the load points for all profiles. This

can be attributed to the damage imparted to the specimen only in the midspan area prior

to testing. Strain increases between the load levels at the prestressing steel layer at the

north load point shown in Figure 4-24 are 157 x 10-6 in/in, 862 x 10-6 in/in, 1146 x 10-6

in/in, and 794 x 10-6 in/in. Strain increases between the load levels at the prestressing

steel layer at midspan shown in Figure 4-25 are 802 x 10-6 in/in, 1121 x 10-6 in/in, 1158 x

10-6 in/in, and 1103 x 10-6 in/in. Strain increases between the load levels at the

prestressing steel layer at the south load point shown in Figure 4-26 are 122 x 10-6 in/in,









1572 x 10-6 in/in, 959 x 10-6 in/in, and 931 x 10-6 in/in. The largest increases for the

strain gauges near the level of prestressing were located at midspan for the first load step

up to 0.25Pmax, the south load point for the next load step up to 0.50Pmax, and the midspan

for the last two load steps up to Pmax. The strain increase pattern shows overall

symmetric strain increases about midspan except between 0.50 and 0.75Pmax. Figure 4-

27 shows the load versus strain data for the strain gauges shown in Figure 3-6a along the

bottom centerline of the test specimen. Similar strain readings were recorded for the four

strain gauges up to the maximum load. Figure 4-27 also shows slightly larger strains at

the maximum load for the strain gauges located on the north end of the girder than for the

strain gauges located on the south end of the girder. The maximum recorded bottom

centerline strain at Pmax was 0.004897 in/in and occurred at 1'-0" north of midspan.

The failure mode of Test Specimen 4 was a combination of FRP rupture and

adhesive failure as shown in Figure 4-28. Before failure, the FRP system had debonded

between the stirrups and was stretched in tension due to the deflection of the girder. At

failure, the FRP had reached its tensile capacity in which the FRP ruptured just before the

stirrup at the south end and debonded past the stirrup at the north end. By debonding past

the stirrup, the longitudinal FRP sheared through the stirrup, which had fibers oriented at

0 and 900 as shown in Figure 4-29. The shear, moment, and midspan deflection

recorded at failure for the repaired girder were 202 kips, 1562 kip-ft, and 1.902 inches,

respectively. The ultimate shear and moment capacity corresponds to restoring 75% and

78%, respectively, of the experimental ultimate capacity of the undamaged girder. The

restored percentages for shear and moment capacity differ due to the supports being

moved out an additional 6" for the test as shown in Figure 3-5. The measured deflection









at Pmax for Test Specimen 4 corresponds to a 68% reduction in the deflection of the

undamaged girder at its maximum load. Loading was discontinued for Test Specimen 4

when the FRP system failed at a load corresponding to less than the ultimate

experimental load of Test Specimen 2, representing the girder in the damaged state. It

can be assumed that Test Specimen 4 would have ultimately failed at approximately the

ultimate load of Test Specimen 2.

Test 5 UF Sprayed Repair Results

The load versus midspan deflection for Test Specimen 5 is shown in Figure 4-30.

The experimental load capacities at cracking and ultimate are 125 kips and 255 kips with

corresponding midspan deflections of 0.53 and 3.16 inches, respectively. The load

deflection relationship, as shown in Figure 4-30 is linear up to the cracking load and

nonlinear between the cracking load and maximum load. Symmetrical deflections

occurred about the midspan of the beam as shown in Figure 4-31, except for a slightly

larger deflection at the north load point at the maximum load. Figure 4-31 also shows

similar increases in deflection between 0 and 25 percent and between 25 and 50 percent

of the maximum load, with slightly larger increases in deflection between 50 and 75

percent, 75 and 90 percent, and 90 and 100 percent of the maximum load.

The strain profiles in Figures 4-32, 4-33, and 4-34 all show an upward movement

of the neutral axis from the initial location of 20.96 inches down from the top of the slab

into the slab at Pmax. The final depths of the neutral axis at the north load point, midspan,

and south load point locations are 16.05", 9.18", and 9.12", respectively, down from the

top of the slab at the maximum load. The significant difference between the neutral axis

depths at the load point locations can be attributed to either a slight imbalance of 6 kips at

the maximum applied load with a higher load at the south load point location, larger









deflections, slipping of prestressing strands, yielding of the strands, compression failure,

or failure of the FRP all on the south end of the girder. The strain profiles in Figures 4-32

and 4-34 for the north and south load point locations show similar strain patterns for all

load increments with slightly larger strains at the south load point. A comparison of the

strains in Figures 4-32, 4-33, and 4-34 show that there were larger increases at midspan

than at the load point locations for all profiles. This can be attributed to the damage

imparted to the specimen only in the midspan area prior to testing. Strain increases

between the load levels at the prestressing steel layer at the north load point shown in

Figure 4-32 are 254 x 10-6 in/in, 2084 x 10-6 in/in, 1451 x 10-6 in/in, and 952 x 10-6 in/in.

Strain increases between the load levels at the prestressing steel layer at midspan shown

in Figure 4-33 are 808 x 10-6 in/in, 2714 x 10-6 in/in, 2487 x 10-6 in/in, and 2814 x 10-6

in/in. Strain increases between the load levels at the prestressing steel layer at the south

load point shown in Figure 4-34 are 209 x 10-6 in/in, 1149 x 10-6 in/in, 1733 x 10-6 in/in,

and 1303 x 10-6 in/in. Comparing the strain increases at the prestressing steel level shows

larger increases in strain at the midspan location than at either of the load point locations

and similar increases in strain at the north and south load point locations. These larger

increases at the midspan location are reasonable since the midspan is within the constant

or maximum moment region and the load points are located outside the constant moment

region. Figure 4-35 shows the load versus strain data for the strain gauges shown in

Figure 3-6a along the bottom centerline of the test specimen. Similar strain readings

were recorded for the two strain gauges at 1'-0" from midspan up to failure and for the

two strain gauges 3'-0" from midspan up to failure. In Figure 4-35, the strain gauge

located 3'-0" from midspan on the south end of the girder is experiencing increased









sensitivity noted by the repeated back and forth motion although the gauge remained

functional throughout the loading. Figure 4-35 also shows larger strains after the

cracking load for the strain gauges 1'-0" from midspan than the strain gauges 3'-0" from

midspan which are located one foot from the end of the damaged area. The maximum

recorded bottom centerline strain at Pmax was 0.009234 in/in and occurred at 1'-0" north

of midspan.

The failure mode of Test Specimen 5 was FRP rupture as shown in Figures 4-36

and 4-37. A crack formed near midspan of the beam rupturing completely through the

FRP applied to the perimeter of the section. The FRP remained attached over the entire

length and perimeter of the girder. The shear, moment, and midspan deflection recorded

at failure for the repaired girder were 255 kips, 1911 kip-ft, and 3.16 inches, respectively.

The ultimate shear and moment capacity corresponds to restoring 95% of the

experimental ultimate capacity of the undamaged girder. The deflection at Pmax for Test

Specimen 5 corresponds to a 47% reduction in the deflection of the undamaged girder at

its maximum load.

Test 6 Edge Composites Repair Results

Test Specimen 6 was not tested at this time.

Evaluation of Experimental Observations

Table 4-1 shows the capacities, midspan deflections, strains, and initial flexural

stiffness at the cracking load and Table 4-2 shows the capacities, midspan deflections,

and strains at the maximum load.

Effect of Fiber Reinforced Polymers on Stiffness

A significant reduction in stiffness occurred when four prestressing strands were

severed in the test specimens to simulate the impact damage. This stiffness loss resulted









in larger deflections at reduced load levels. The maximum load and subsequent failure

also occurred at reduced load levels. The experimental initial flexural stiffness of the test

specimens was determined by dividing the load at cracking by the deflection at cracking.

The initial flexural stiffness calculated for Test Specimen 1 was 262 kips/in while Test

Specimen 2 had an initial flexural stiffness of 208 kips/in as shown in Table 4-1.

Therefore, a 21% reduction in initial flexural stiffness was simulated by the severing of

four prestressing strands as shown in Table 4-1.

The application of FRP materials to the tension face of a girder provides resistance

to deflection through the high tensile strength of the FRP materials in comparison to

concrete or mild steel. The resulting amount of resistance or stiffness increase is

dependent on FRP laminate properties, FRP location on the girder, thickness of the FRP

laminate, and an effective bond. Test Specimen 3, repaired with carbon fibers and an

epoxy adhesive had an initial stiffness that was only of 79.5% of Test Specimen 1; Test

Specimen 4, repaired with carbon fibers with a polyurethane adhesive had an initial

stiffness that was 89.4% of Test Specimen 1; and Test Specimen 5, repaired with glass

fibers with a polyester adhesive had a comparable initial stiffness that was 89.5% of Test

Specimen 1 as shown in Table 4-1. The resulting initial flexural stiffness increases of

Test Specimens 3, 4, and 5 over Test Specimen 2 were 0%, 12.8%, and 12.9%,

respectively, as shown in Table 4-1.

At the cracking load the experimentally measured midspan deflections of Test

Specimens 2, 3, 4, and 5 varied in comparison to the measured midspan deflection of Test

Specimen 1. The measured midspan deflection at cracking load for Test Specimen 1 was

0.50" while Test Specimen 2 had a midspan deflection at cracking load of 0.54" as shown









in Table 4-1. Therefore, a 7.6% increase in midspan deflection at cracking load was

simulated by the severing of four prestressing strands. Test Specimen 3, with a midspan

deflection at cracking load equal to 0.60", had a midspan deflection at cracking load 20%

greater than Test Specimen 1; Test Specimen 4, with a midspan deflection at cracking

load equal to 0.47", had a midspan deflection at cracking load 5.3% less than Test

Specimen 1; and Test Specimen 5, with a midspan deflection at cracking load of 0.53",

had a midspan deflection at cracking load 6.6% greater than Test Specimen 1 as shown in

Table 4-1. The measured midspan deflection at maximum load for Test Specimen 1 was

5.95" while Test Specimen 2 had a midspan deflection at cracking load of 6.42" as shown

in Table 4-2. Therefore, a 8% increase in midspan deflection at maximum load was

simulated by the severing of four prestressing strands. Test Specimen 3, with a midspan

deflection at maximum load equal to 3.15", had a midspan deflection at maximum load

47% less than Test Specimen 1; Test Specimen 4, with a midspan deflection at maximum

load equal to 1.90", had a midspan deflection at maximum load 68% less than Test

Specimen 1; and Test Specimen 5, with a midspan deflection at maximum load of 3.16",

had a midspan deflection at maximum load 47% less than Test Specimen 1 as shown in

Table 4-2.

Effect of Fiber Reinforced Polymers on Capacity

A significant reduction in capacity occurred when four prestressing strands were

severed in the test specimens to simulate impact damage. This loss in capacity resulted in

failures at reduced load levels. The shear and moment capacity of Test Specimen 1 at the

cracking load was 131 kips and 984 kip-ft while the shear and moment capacity of Test

Specimen 2 at the cracking load was 112 kips and 839 kip-ft as shown in Table 4-1.

Therefore, a 15% reduction in the shear and moment capacity at the cracking load was









simulated by the severing of four prestressing strands. The experimental capacity at the

cracking load for Test Specimen 3 was 125 kips which was equal to 96% of the capacity

of Test Specimen 1, and equal to 112% gain over the capacity of Test Specimen 2 as

shown in Table 4-1. The experimental capacity at the cracking load for Test Specimen 4

was 111 kips which was equal to 85% of the capacity of Test Specimen 1, and equal to

99% of the capacity of Test Specimen 2 as shown in Table 4-1. The experimental

capacity at the cracking load for Test Specimen 5 was 125 kips which was equal to 95%

of the capacity of Test Specimen 1, and equal to 112% gain over the capacity of Test

Specimen 2 as shown in Table 4-1. The shear and moment capacity of Test Specimen 1

at the maximum load was 268 kips and 2012 kip-ft, respectively. While the shear and

moment capacity of Test Specimen 2 at the maximum load was 210 kips and 1576 kip-ft,

respectively as shown in Table 4-2. Therefore, a 22% reduction in the shear and moment

capacity at the maximum load was simulated by the severing of four prestressing strands.

The experimental capacity at the maximum load for Test Specimen 3 was 244 kips which

was equal to 91% of the capacity of Test Specimen 1, and equal to 116% over the

capacity of Test Specimen 2 as shown in Table 4-2. The experimental capacity at the

maximum load for Test Specimen 4 was 202 kips which was equal to 75% of the capacity

of Test Specimen 1, and equal to only 96% of the capacity of Test Specimen 2 as shown

in Table 4-2. The experimental capacity at the maximum load for Test Specimen 5 was

255 kips which was equal to 95% of the capacity of Test Specimen 1, and equal to a

121% over the capacity of Test Specimen 2 as shown in Table 4-2. The amount of

capacity increase that the FRP can provide to a structural member is dependent on FRP