FATIGUE EVALUATION OF TWO VARIABLE MESSAGE SIGN STRUCTURES. Final Report. T. Weston McLean Jong Sup Park J. Michael Stallings.

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1 FATIGUE EVALUATION OF TWO VARIABLE MESSAGE SIGN STRUCTURES Final Report by T. Weston McLean Jong Sup Park J. Michael Stallings sponsored by The Alabama Department of Transportation Montgomery, Alabama July 24

2 ABSTRACT Transportation departments of most states do not perform routine inspections of sign structures. This makes it very important that fatigue cracking of sign structures be avoided. In recent years, Variable Message Sign (VMS) panels have become increasingly popular. The geometry of these signs sometimes results in significant cyclic loading of the support structure due to wind gusts. These wind gusts can result from passing trucks or from natural wind. This report covers analytical and experimental investigations of VMS structures in Birmingham and Mobile, Alabama. Commercially available software GTSTRUDL Is used to perform space frame structural analyses of these welded tubular structures. Measurements from a static load test of the Birmingham structure are presented and compared to structural analysis results. These comparisons indicate generally good agreement. Fatigue evaluations are performed using stress ranges from field measurements and from structural analyses. Applied loadings in the structural analyses are the natural wind gust and truck-induced wind gust fatigue design loadings of the Standard Specifications for Structural Supports for Highway Signs, Luminaires and Traffic Signals, 4 th Edition published in 21 (AASHTO 21). The structures studied were designed prior to publication of AASHTO 21. Based on structural analyses using the fatigue design loadings from that specification, the structures studied are found not to have infinite fatigue life. Using stress ranges from field measurements, fatigue cracking is predicted in five members of the Birmingham structure. But, only one of these five members has a predicted fatigue life of less i

3 than the 5-year design life. Using stress ranges from field measurements, the Mobile structure is predicted to have infinite life. Based on the limited measurements made in this project, the fatigue design loadings of AASHTO (21) appear to be conservative, but not overly conservative. ii

4 TABLE OF CONTENTS TABLES FIGURES... vi... viii CHAPTER 1. INTRODUCTION Background Research Sites Description of Sign Support Structure Birmingham Structure Mobile Structure Previous Research Project Scope Project Objectives... 7 CHAPTER 2. LITERATURE REVIEW Introduction Wind Loading Relevant to VMS Support Structures Natural Wind Gusts Truck-Induced Wind Gusts Fatigue Categories of Connection Details Background on Fatigue Resistance Fatigue Life Infinite Life Design... 2 CHAPTER 3. STRUCTURAL ANALYSES AND RESULTS Introduction iii

5 3.2 Space Frame Models Static Analyses Pull Down Loading Condition Birmingham Structure Natural Wind Loads Truck-Induced Gusts Loading Dynamic Analyses Birmingham Structure Mobile Structure Conclusions CHAPTER 4. FIELD TESTS AND FATIGUE EVALUATION Introduction Instrumentation and Data Acquisition Strain Gage Locations Birmingham Structure Mobile Structure Strain Gages Wind Anemometer Data Acquisition System Field Tests Response to Truck-Induced Gusts Birmingham Structure Mobile Structure Static Pull Down Test Birmingham Structure iv

6 Testing Procedure Data Reduction Results Long-Term Monitoring Birmingham Structure Mobile Structure Stress Range Occurring at 1 in 1, Cycles Effective Stress Range Fatigue Evaluation Finite Life Check Fatigue Life Evaluation Conclusions CHAPTER 5. CONCLUSIONS Project Objectives Conclusions REFERENCES APPENDIX A APPENDIX B APPENDIX C v

7 TABLES Table 2.1 AASHTO (21) Detail Category A-Values Table 3.1 Member Properties Birmingham Structure Table 3.2 Member Properties Mobile Structure Table 3.3 Pull Down Loading Cases Table 3.4 Deflection due to Pull Down Loading Cases Table 3.5 Structural Analysis Results for Trial Table 3.6 Natural Wind Gust Loads Applied to VMS Panels Table 3.7 Stresses at Strain Gage Locations due to Natural Wind Gust Birmingham Structure Table 3.8 Stresses of Strain Gage Points Due to Natural Wind Gust Mobile Structure Table 3.9 Truck-Induced Gust Loadings Applied to VMS Panels Table 3.1 Truck-Induced Loadings Applied to Truss Members Behind VMS Table 3.11 Stresses at Strain Gage Locations due to Truck-Induced Gust Birmingham Structure Table 3.12 Stresses of Strain Gage Points Due to Truck-Induced Gust 4Mobile Structure... 4 Table 3.13 Modal Frequencies Birmingham Structure... 4 Table 3.14 Modal Frequencies Mobile Structure Table 4.1 Comparisons of ulated and Measured Truck-Induced Stress Ranges Birmingham Structure Table 4.2 Comparisons of ulated and Measured Truck-Induced Stress Ranges Mobile Structure Table 4.3 Static Pull Down Load Increments for Each Trial Table 4.4 Peak Load Comparisons at Main Chord Members vi

8 Table 4.5 Field Measured Stress Cycle Summary Birmingham Structure Table 4.6 Field Measured Stress Cycle Summary - Mobile Structure Table 4.7 Comparisons of ulated and Measured Limit-State Stress Ranges - Birmingham Structure... 8 Table 4.8 Comparisons of ulated and Measured Limit-State Stress Ranges - Mobile Structure Table 4.9 Estimated Yearly Stress Range Histograms - Birmingham Structure Table 4.1 Estimated Yearly Stress Range Histograms - Mobile Structure Table 4.11 Summary of Measured Effective Stress Range with Comparisons to Measured Maximum and Limit-State Stress Ranges Birmingham Structure. 84 Table 4.12 Summary of Measured Effective Stress Range with Comparisons to Measured Maximum and Limit-State Stress Ranges Mobile Structure Table 4.13 Fatigue Life Estimates for all Gage Locations Birmingham Structure 85 Table 4.14 Fatigue Life Estimates for all Gage Locations Mobile Structure Table B.1 Measured Results for Trial 1 of Static Load Pull Down Test Table B.2 Measured Results for Trial 2 of Static Load Pull Down Test Table B.3 Measured Results for Trial 3 of Static Load Pull Down Test Table B.4 Measured Results for Trial 4 of Static Load Pull Down Test Table B.5 Measured Results for Trial 5 of Static Load Pull Down Test Table B.6 Structural Analysis Results for Trial Table B.7 Structural Analysis Results for Trial Table B.8 Structural Analysis Results for Trial Table B.9 Structural Analysis Results for Trial vii

9 FIGURES Figure 1.1. Southbound View of VMS Structure in Birmingham, AL Figure 1.2. Eastbound View of VMS Structure in Mobile, AL Figure 1.3. Truss to Vertical Support Connection of Birmingham Structure... 9 Figure 1.4. Vertical Support Foundation of Birmingham Structure... 9 Figure 1.5. Truss to Vertical Support Connection of Mobile Structure... 1 Figure 1.6. Vertical Support Foundation of Mobile Structure... 1 Figure 1.7. Typical Splice Connection Figure 2.1. S-N Curves for Detail Categories A through ET ( from AWS 2) Figure 2.2. Example of Detail Category E and ET (AASHTO 21) Figure 3.1. Model of VMS Support Structure on US-28 in Birmingham Figure 3.2. Model of VMS Support Structure on I-1 in Mobile Figure 3.3. Pull Down Loading Condition of Birmingham Support Structure Figure 3.4. Typical Deformed Shape of Birmingham Structure Due to Pull Down Loading Condition Figure 3.5. Static Loading Condition of Birmingham VMS Due to Natural Wind Gust Figure 3.6. Static Loading Condition of Birmingham Support Structure Due to Natural Wind Gust Figure 3.7. Static Loading Condition of Mobile VMS Due to Natural Wind Gust Figure 3.8. Static Loading Condition of Mobile Support Structure Due to Natural Wind Gust Figure 3.9. Static Loading Condition of Birmingham VMS Due to Truck-Induced Gust Figure 3.1. Static Loading Condition of Birmingham Support Structure Due to Truck-Induced Gust viii

10 Figure Static Loading Condition of Mobile VMS Due to Truck-Induced Gust.. 49 Figure Static Loading Condition of Mobile Support Structure Due to Truck-Induced Gusts Figure First Mode Shape of Birmingham Support Structure... 5 Figure Second Mode Shape of Birmingham Support Structure... 5 Figure Third Mode Shape of Birmingham Support Structure Figure Fourth Mode Shape of Birmingham Support Structure Figure First Mode Shape of Mobile Support Structure Figure Second Mode Shape of Mobile Support Structure Figure Third Mode Shape of Mobile Support Structure Figure 3.2. Fourth Mode Shape of Mobile Support Structure Figure 4.1. Diagram of Strain Gage Locations for Right Half of Birmingham Structure Figure 4.2. Detail of Strain Gage Placement on Chords and Diagonals Figure 4.3. WT Section Strain Gage Placement on Birmingham Structure Figure 4.4. Diagram of Strain Gage Locations on Mobile Structure Figure 4.5. Truck-Induced Vibration Response of Birmingham Structure from Strain Gage H Figure 4.6. Truck-Induced Vibration Response of Birmingham Structure fromr Strain Gage FT Figure 4.7. Truck-Induced Vibration Response of Birmingham Structure from Strain Gage T Figure 4.8. Truck-Induced Vibration Response of Mobile Structure from Strain Gage V Figure 4.9. Pull Down Test Setup Figure 4.1. Load Cell and Pulley ix

11 Figure Deflectometer Figure Recording of Data During Pull Down Test Figure Typical Strain Output for Pull Down Loading Test (Trials 2 through 5) 97 Figure Illustration of Data Reduction for Each Trial Figure 4.15 ulated and Measured Load Versus Stress for Gage BB Figure 4.16 ulated and Measured Load Versus Stress for Gage BB Figure 4.17 ulated and Measured Load Versus Stress for Gage BB Figure 4.18 ulated and Measured Load Versus Stress for Gage FT Figure 4.19 ulated and Measured Load Versus Stress for Gage FT Figure 4.2 ulated and Measured Load Versus Stress for Gage FT Figure 4.21 ulated and Measured Load Versus Stress for Gage FT Figure 4.22 ulated and Measured Load Versus. Stress for Gage BT Figure 4.23 ulated and Measured Load Versus Stress for Gage BT Figure 4.24 ulated and Measured Load Versus Stress for Gage BT Figure 4.25 ulated and Measured Load Versus Stress for Gage BT Figure 4.26 ulated and Measured Load Versus Stress for Gage S Figure 4.27 ulated and Measured Load Versus Stress for Gage S Figure 4.28 ulated and Measured Load Versus Stress for Gage V Figure 4.29 ulated and Measured Load Versus Stress for Gage V Figure 4.3 ulated and Measured Load Versus Stress for Gage H Figure 4.31 ulated and Measured Load Versus Stress for Gage H Figure 4.32 ulated and Measured Load Versus Stress for Gage H Figure 4.33 ulated and Measured Load Versus Stress for Gage H x

12 Figure 4.34 ulated and Measured Load Versus Stress for Gage H Figure 4.35 ulated and Measured Load Versus Stress for Gage V Figure 4.36 ulated and Measured Load Versus Stress for Gage V Figure 4.37 ulated and Measured Load Versus Stress for Gage T Figure 4.38 ulated and Measured Load Versus Stress for Gage T Figure 4.39 ulated and Measured Load Versus Deflection for Bottom Chords123 Figure 4.4 Wind Anemometer Output During High Wind Event Figure 4.41 Effective Stress Ranges for One Day Periods for Front Top Chord on Birmingham Structure Figure 4.42 Effective Stress Ranges for One Day Periods for Gage H1 on Birmingham Structure Figure 4.43 Effective Stress Ranges for One Day Periods for Gage H1 and H2 on Mobile Structure Figure C.1. ulated and Measured Load vs. Stress for Gage S Figure C.2. ulated and Measured Load vs. Stress for Gage S Figure C.3. ulated and Measured Load vs. Stress for Gage V Figure C.4. ulated and Measured Load vs. Stress for Gage V Figure C.5. ulated and Measured Load vs. Stress for Gage V Figure C.6. ulated and Measured Load vs. Stress for Gage V Figure C.7. ulated and Measured Load vs. Stress for Gage H Figure C.8. ulated and Measured Load vs. Stress for Gage H Figure C.9. ulated and Measured Load vs. Stress for Gage H Figure C.1. ulated and Measured Load vs. Stress for Gage H Figure C.11. ulated and Measured Load vs. Stress for Gage H Figure C.12. ulated and Measured Load vs. Stress for Gage FT xi

13 Figure C.13. ulated and Measured Load vs. Stress for Gage FT Figure C.14. ulated and Measured Load vs. Stress for Gage FT Figure C.15. ulated and Measured Load vs. Stress for Gage FT Figure C.16. ulated and Measured Load vs. Stress for Gage BT Figure C.17. ulated and Measured Load vs. Stress for Gage BT Figure C.18. ulated and Measured Load vs. Stress for Gage BT Figure C.19. ulated and Measured Load vs. Stress for Gage BT Figure C.2. ulated and Measured Load vs. Stress for Gage BB Figure C.21. ulated and Measured Load vs. Stress for Gage BB Figure C.22. ulated and Measured Load vs. Stress for Gage BB Figure C.23. ulated and Measured Load vs. Stress for Gage T Figure C.24. ulated and Measured Load vs. Stress for Gage T Figure C.25. ulated and Measured Load vs. Deflection for Front and Back Chords xii

14 CHAPTER 1. INTRODUCTION 1.1 BACKGROUND Variable Message Signs (VMS) have become a new trend in roadway signs. These signs offer an increase in traffic safety by their ability to relay messages to motorists for warnings of hazards ahead such as fog, traffic congestion, accidents, and lane closings. These signs can also be equipped to monitor traffic speed and flow. VMS panels are generally installed over traffic lanes on heavily traveled roadways such as interstates and major highways. These sign panels are relatively heavy and larger in size when compared to a typical flat panel sign board. The increase in size and weight results in greater wind load and inertial effects. Natural winds are a major cause of horizontal vibrations in these structures. Trucks passing beneath these signs create wind gusts which produces loading on the VMS both in the upward direction and in the direction of traffic flow. As a result, these VMS structures are subjected to large number of loading cycles which can result in fatigue cracking. The AASHTO Specifications for Structural Supports for Highway Signs, Luminaires, and Traffic Signals (1994) offers little guidance in the design for fatigue. Until the most recent version (AASHTO 21) of these specifications became available, designers did not have sufficient direction for fatigue related design of signs, signals and luminaire support structures. As a result, some cantilevered sign and signal support structures across the country have exhibited excessive displacement due to wind-induced vibration and have failed as a result of fatigue cracking. 1

15 1.2 RESEARCH SITES VMS panels have been installed on several major highways and interstates in Birmingham and Mobile, Alabama. The Alabama Department of Transportation (ALDOT) officials have observed significant vibration of a bridge type VMS support structure located on U.S. Highway 28 just south of the U.S. Highway 28/I-459 interchange in Birmingham. Those observations were a primary motivating factor for undertaking the work reported here. This particular structure has one of the longer span lengths for bridge type support structures, spanning 145 ft across all lanes of traffic. Figure 1.1 is a southbound view of the Birmingham VMS structure. Another bridge type VMS structure in Mobile, Alabama is included in this study. This structure is located on I-1 just west of the tunnel. It spans 89 ft over the all eastbound lanes of traffic. It was chosen for this study because of the high wind speeds common for this location. An eastbound view of the Mobile structure is shown in Figure DESCRIPTION OF SIGN SUPPORT STRUCTURES Birmingham Structure Figures 1.3 and 1.4 show the truss-to-vertical support connection and the vertical support-to-foundation connection. The VMS support structure is a four chord box pipe truss with fillet welded tubular members as the web of the truss. The truss spans over all six lanes of traffic; 3 northbound lanes and 3 southbound lanes with a median between them. The truss consists of four sections connected at each of the four chords by transverse plates bolted together and is geometrically similar about 2

16 the midspan connection. The horizontal truss is supported at each end by a vertical two post truss support. The two bottom chords of the truss rest on a horizontal member made from a WT-section that spans between two vertical support posts. The bottom chords of the truss are bolted to the WT-section using U-bolts while the two top chords are connected to either vertical post also using U-bolts. The VMS is centered over the three northbound lanes 19 ft above the roadway and is attached to W6x9 hangers bolted to the front two chords with U-bolts Mobile Structure Figures 1.5 and 1.6 show the truss-to-vertical support connection and foundation of the right support. The truss consists of 3 equal sections connected at the four chords by transverse plates bolted together. All connections are similar to the structure in Birmingham with the exception of the truss-to-vertical support connection. As shown in Figure 1.5, the vertical support posts do not extend up to the top two chords of the truss so there is no connection between the top two chords and the vertical supports. The two bottom chords rest on the WT-section and are bolted down using U-bolts. The VMS is centered over the four northbound lanes of traffic. Figure 1.7 illustrates connections typical for both Birmingham and Mobile structures. For convenience and ease of referencing different locations on both structures, the vertical side (if one is facing the VMS) will be denoted throughout this report as the front plane or side of the structures. The notations back, top, and bottom will be used to describe the other three planes of the box truss structure relative to the front side. The notations right and left will be used in the same respect 3

17 (i.e. the right support is on the right if one is facing the VMS). This designation scheme will also be used later when referencing a particular strain gage location. 1.4 PREVIOUS RESEARCH Numerous research projects have been conducted on cantilever type support structures; however, very few investigations of bridge type support structures have been performed. Some state transportation departments have had problems with cantilever VMS structures such as excessive vibration and fatigue. Failures of cantilevered support structures for VMS have occurred in Virginia and California. There was a collapse in 1993 of a VMS in Virginia that was believed to be the result of fatigue cracking due to truck-induced wind gusts. Truck-induced wind gusts were also suspected to be the cause of excessive vibration of VMS structures in New Jersey and Florida (Dexter and Ricker 22). Variable Message Signs have also been installed on support structures originally intended to support only flat panel signs that weigh much less than VMS. This was done without making calculations of the effect of the additional load and area of the VMS. The additional mass was thought to cause problems; however, engineers have discovered that the mass only adds to the dead load and can be mitigated by increasing the stiffness of the supporting structure. Therefore, the additional exposed area of the sign is believed to cause the fatigue problems (Dexter and Ricker 22). 4

18 1.5 PROJECT SCOPE Originally the goal of the project reported here was to investigate the feasibility of mounting variable message sign (VMS) panels onto existing structures that were designed to support flat panel signs. A conclusion was reached relatively quickly that retrofitting existing structures would not be a desirable long-term practice for the Alabama Department of Transportation (ALDOT). Obvious vibration of a new VMS structure on Highway 28 in Birmingham, Alabama was reported by ALDOT Division personnel. An evaluation of the significance of these vibrations became a primary goal of the project reported here. The potential failure mode investigated was fatigue cracking at the welded connections of the tubular members of the structure. An analytical and experimental investigation began by performing preliminary structural analyses to determine locations of high stress ranges due to the fatigue design loads. The results were used in planning the experimental instrumentation. These members were then instrumented with strain gages and monitored for an extended period to evaluate the structural response to natural wind and truckinduced gusts. The nominal stress ranges were measured at these strain gage locations to determine if the stress ranges were large enough to eventually result in fatigue cracking. The number of loading cycles and effective stress ranges were also needed to perform a remaining fatigue life evaluation. Chapter 4 presents the instrumentation and field testing procedures. Significant changes from the previous design specifications were incorporated into the AASHTO Specifications for Structural Supports for Highway Signs, 5

19 Luminaires and Traffic Signals in 21. These revisions include fatigue design provisions that define fatigue design loadings from natural wind gusts and truckinduced wind gusts. The fatigue design provisions were developed for cantilever sign structures, but can be applied to bridge type structures such as the Birmingham structure. A second primary goal of the project reported here was to investigate the applicability of the AASHTO (21) fatigue design provisions to bridge type structures. The Birmingham structure and a VMS sign structure in Mobile were used as example structures for application of the AASHTO (21) provisions. Both of these structures were designed before AASHTO (21) was published. Instrumentation installed to monitor stress cycles due to wind gusts provided an opportunity to assess the accuracy of structural analysis type of these structure. This was achieved by performing a static pull down load test on the Birmingham structure where strains and displacements were measured under a known applied load. Comparisons to the analytical results were then made to assess the accuracy of the analyses. The long-term monitoring and static load test results along with comparisons to the analytical results are presented in Chapter 4. The fatigue design provisions of AASHTO (21) were used in an evaluation of the Birmingham structure. This evaluation is summarized in Appendix A. GTSTRUDL input files for the structural analyses are on the disk provided with this report. 6

20 1.6 PROJECT OBJECTIVES Specific objectives of the project are as follows: (1) Determine whether the noticeable vibration of the Birmingham structure is an indication of an unsafe condition. (2) Determine the fatigue stress ranges being applied to the Birmingham and Mobile structures and estimate their fatigue life. (3) Determine whether these structures satisfy the fatigue design provisions of AASHTO (21). (4) Investigate the accuracy of a standard space frame analysis of a tubular sign structure. (5) Determine whether the fatigue design wind gust loadings of AASHTO (21) appear reasonable. Conclusions regarding these objectives are stated in Chapter 5. 7

21 Figure 1.1. Southbound View of VMS Structure in Birmingham, AL. Figure 1.2. Eastbound View of VMS Structure in Mobile, AL. 8

22 Figure 1.3. Truss to Vertical Support Connection of Birmingham Structure Figure 1.4. Vertical Support Foundation of Birmingham Structure 9

23 Figure 1.5. Truss to Vertical Support Connection of Mobile Structure Figure 1.6. Vertical Support Foundation of Mobile Structure 1

24 Figure 1.7. Typical Splice Connection 11

25 CHAPTER 2. LITERATURE REVIEW 2.1 INTRODUCTION The VMS structures were designed based on the AASHTO Standard Specifications for Structural Supports for Highway Signs, Luminaires and Traffic Signals (1994). However, the provisions outlined in this 1994 edition of the specification are vague and insufficient with respect to the design of structures for vibration and fatigue. Furthermore, the commentary to the specifications does not contain adequate guidance for the application of these provisions. Recent research performed under the NCHRP Project 1-38 by Dexter, Kaczinski, and Van Dien (1996) has led to an entire fatigue design chapter (Section 11) in AASHTO (21). According to AASHTO (21), sign, signal, and luminaire support structures are designed to resist fatigue caused by each of the applicable equivalent static wind load effects described below. Stress ranges due to these loads on all components, mechanical fasteners, and weld details are designed to satisfy the requirements of their respective detail categories with the constant-amplitude fatigue thresholds given in the specifications (AASHTO 21). Accurate load spectra and life prediction estimates for defining fatigue loadings are typically not available for designers. Moreover, the assessment of stress histories and the corresponding lifetime wind loading histograms are practically impossible. As a result, the design of support structures for a finite life becomes impractical; therefore, an infinite-life fatigue design approach is recommended. To provide infinite life, fatigue critical details are designed so that the nominal applied stress ranges are below the constant amplitude fatigue limit (CAFL) appropriate for the particular detail (AASHTO 21). 12

26 Table 3-3 in AASHTO (21) recommends the minimum design life for different types of support structures. Luminaires exceeding 15 m (49.2 ft) in height and overhead sign structures (cantilevered and bridge supports) are designed for 5 years. Smaller luminaires and traffic signal supports are designed for a minimum life of 25 years. 2.2 WIND LOADING RELEVANT TO VMS SUPPORT STRUCTURES There are four types of wind loading phenomena that are critical with respect to the design for vibration and fatigue of sign, signal, and luminaire support structures. These are 1) galloping; 2) vortex shedding; 3) natural wind gusts; and 4) truck induced wind gusts. Cantilevered VMS support structures are susceptible to galloping, natural wind gusts, and truck-induced wind gusts. Bridge type VMS support structures are not expected to be susceptible to galloping due to the torsional rigidity of the sign bridge. However, bridge type support structures may be susceptible to natural wind gusts and truck-induced wind gusts. The application of complex dynamic wind loads such as those due to the phenomena listed above is simplified in AASHTO (21) through the use of equivalent static loads. Although these fatigue loadings were developed for use on cantilevered support structures, the commentary to Section 11.5 of AASHTO (21) specifies that these loads can be applied to bridge type structures until more appropriate loadings are developed. These equivalent static loads are considered limit-state fatigue loads. Fatigue limitstate stress ranges induced by dynamic wind loads are estimated with static loads that would create similar stress responses. Therefore a simple static analysis is 13

27 conducted by designers and the resulting stresses represent stress ranges. These stress ranges are limited to the appropriate CAFL to ensure infinite life of the structure (AASHTO 21) Natural Wind Gusts Natural wind gusts are inherently variable in the direction and velocity of the air flow. Natural winds are characterized by two components; a mean wind velocity component and a fluctuating component. This fluctuating component of the wind induces fluctuating pressures on the various structural members of a wind-loaded structure which may result in vibration of the structure. The response of a structure to natural wind gusts is variable and randomly distributed (Dexter, Kaczinski, and Van Dien 1996). Therefore, the destructive response of resonance is not likely to occur. The magnitudes of the variable stress ranges induced in the connection details of a structure may, in some cases, be such that fatigue cracking is possible due to the long term cumulative effect of these natural wind gusts (Dexter, Kaczinski, and Van Dien 1996). Using AASHTO (21), the natural wind gust pressure is applied horizontally to the projected frontal area of all surfaces, including the structural members as well as the sign and signal attachments (Dexter et al. 22). The natural wind pressure is calculated as: P = 5.2C I ( psf ) (2.1) NW d F where: C d = drag coefficient I F = importance factor 14

28 The above equation is based on the.1 percent exceedence for a yearly mean wind velocity of 5 m/s (11.2 mph). According to AASHTO (21), this is a reasonable upper-bound of yearly mean wind velocities for most locations Truck-Induced Wind Gusts The passage of trucks under support structures tends to induce gust loads on the frontal area and underside of the members and attachments. The magnitude of the pressures caused by trucks in the horizontal direction is much smaller than that of natural wind gusts. Dexter and Johns (1998) concluded for the purpose of fatigue design, truck induced wind loads normal to the front of the sign are not critical. AASHTO (21) specifies that truck-induced gust loads should be applied only in the vertical direction. The truck-induced pressure is calculated as: P = 18.8C I ( psf ) (2.2) TG d F Support structures with VMS panels attached are much more susceptible to truck-induced wind gusts in the vertical direction. The relatively large thickness in the direction of traffic flow creates a large horizontal surface area. Recent vibration problems with sign structures that have large projected areas in the horizontal plane, such as VMS enclosures, have focused attention on the vertical gust pressures created by the passage of trucks beneath the sign (AASHTO 21). The truck-gust pressure is proportional to the speed of the trucks; therefore, signs located on major highways are more susceptible than those where trucks are traveling at lower speeds. The wind velocity should be reduced based on roadway speed because not all structures would be subjected to 65 mph speed zones (Fouad 15

29 et al. 1998). There is also a vertical gradient for the truck-induced gusts, so the clearance between the truck and the structure is critical to the susceptibility to truckinduced vibration. The truck gust loads essentially deminish to zero at a height of 1 m (32.8 ft) above the roadway (Dexter et al. 22). 2.3 FATIGUE CATEGORIES OF CONNECTION DETAILS The current specifications (AASHTO 21) contain provisions which specify that cantilevered support structures should be designed using an infinite life design approach in accordance with the Standard Specifications for Highway Bridges (AASHTO 1996). The following section describes the categorization of typical support structure connection details and describes the fatigue strength of the welded details. The details are categorized using the fatigue design curves in the AASHTO (1996) bridge specifications and/or the Structural Welding Code-Steel (AWS 2). Fatigue is a complex phenomenon governed by factors which are highly variable and difficult to quantify. Previous research has indicated that the fatigue resistance of welded details can be characterized by two main parameters: (1) nominal stress range and (2) notch severity. The notch severity describes the severity of the stress concentration associated with a welded detail. It includes the effects of the global stress concentration associated with the configuration of the detail and the effects of local stress concentration due to the geometry of the weld and the existence of any weld discontinuities. Previous research has also indicated that the yield strength of the material, the mean stress levels, and operating 16

30 temperatures have little influence on the fatigue resistance of full scale, welded structural connections (Dexter, Kaczinski, and Van Dien 1996). The provisions of the bridge specifications (AASHTO 1996) for the design of structures for fatigue are based upon a nominal stress approach in which details are grouped into categories according to their relative fatigue resistance. Each category corresponds to an S-N curve. An S-N curve defines a linear relationship between number of cycles to failure and applied constant amplitude stress range. This specification contains seven design S-N curves labeled A though E in order of decreasing fatigue strength (see Figure 2.1). Fillet-welded tube to transverse plate connections are classified as a Category E detail. The classification for fillet-welded tubular joints is taken from (AWS 2). Category ET was established for this type of connection where the main member radius to thickness ratios (r/t) is below 24. Category E and ET details (see Figure 2.2) are common in truss support structures and are pertinent to this research project. 2.4 BACKGROUND ON FATIGUE RESISTANCE Fatigue is a cumulative damage process caused by the repeated application of loads that results in cracking and possibly rupture of components. Fatigue failures can occur at applied stress levels much lower than the tensile strength of a given material. The fatigue life of a component may also be considered in two phases. The portion of life which occurs before the appearance of a visible crack is known as the crack initiation life, and the remaining life from crack initiation to unstable crack growth is known as the propagation life (South 1994). 17

31 Fatigue cracks form and propagate from weld discontinuities and/or stress concentrations when a structural member is subjected to significant cyclic live loads. When structural members are tested, the loading is typically described in relation to the nominal stress in the loaded member apart from the weld detail. The local stress concentration effect associated with the shape of the weld is considered part of the fatigue resistance of the detail (Dexter, Kaczinski, and Van Dien 1996). 2.5 FATIGUE LIFE Fatigue tests are performed at a number of different stress ranges, and the data are generally plotted with the logarithm of the nominal stress range as the ordinate and the logarithm of the number of cycles to failure as the abscissa. The relationship used to represent the lower bound to this data is referred to as an S-N curve. An S-N curve is a power equation of the form: or N = AS m logn = log A m logs (2.3) (2.4) where N is the number of cycles to failure, A is a constant that depends on the detail category, S is the applied constant amplitude stress range, and m is the inverse of the slope of the S-N curve. Experiments performed by Peronto (23) revealed that the values of A for detail category E and ET were very close to the values recommended in AWS (2) and used by AASHTO (21). Table 2.1 shows the AASHTO detail category values of A and constant amplitude fatigue limits for the AASHTO (21) detail categories relevant to this report 18

32 The AASHTO and AWS S-N curves for fatigue design are based on constantamplitude loading tests with lives typically less than two million cycles. These curves are used to make fatigue life predictions of structures subjected to numbers of cycles greater than two million, but with smaller, variable stress amplitudes. A common method used to predict fatigue lives of details subjected to variable amplitude loadings is to use Miner s Rule to convert a variable amplitude stress history to an equivalent constant-amplitude stress range. Miner s Rule is a linear damage accumulation method developed by Miner in 1945 (Fisher, Kulak, and Smith 1997). It simply assumes that the damage fraction that results from any particular stress range level is a linear function of the number of cycles that takes place at that stress range. If the constant m of the S-N curve defined by Equation 2.4 is equal to 3, then the relative damage of a stress range is proportional to the cube of the stress range. An effective, or equivalent, constant-amplitude stress range, S Re that would cause an equivalent amount of fatigue damage as the variable stress ranges at a given number of cycles, can be defined as follows when m = 3. Re 3 [ ] 1/ 3 S γ i S = Ri (2.5) where S Ri = magnitude of stress range i, γ i = fraction of cycles at stress range i to total cycles. Constant amplitude fatigue test results typically reveal a constant-amplitude fatigue limit (CAFL). This limit corresponds to the stress range at constant-amplitude loading below which the fatigue life appears to be infinite. Infinite life is often interpreted to be a fatigue life greater than the typical service life of the structure, 19

33 usually 5 years in the case of cantilever signal, sign, and luminaire support structures. Infinite life only occurs if practically all stress ranges due to constantamplitude or variable amplitude loading are below the CAFL. In the case where some load cycles exceed the fatigue limit and some are below the limit, the lower bound of test data appears to follow the extrapolation of the S-N curve below the CAFL. A discussion of what fraction of the applied load cycles that must be above the CAFL to result in finite life is discussed in the next section. To calculate an effective stress range using Equation 2.5 the number of cycles at each applied stress range is required. Theoretically, a complete histogram of the loading events for all events during the life of the structure is needed. This is practically impossible to obtain. Hence, field measurements are typically made for a period of time sufficient to represent the entire service life of the structure. 2.6 INFINITE LIFE DESIGN In the NCHRP Report 354 (Fisher et al. 1993), the infinite-life approach was studied using full-scale, long-life variable amplitude tests on fatigue critical details. This study concluded that failure could still occur if.5 percent or more of the stress ranges exceeded the CAFL, and infinite life resulted when.1 percent or fewer of the cycles exceeded the CAFL. The infinite-life variable amplitude fatigue design approach recommended in NCHRP Report 354 suggests that the response of the structure which has a probability of exceedence of only.1 percent must be estimated. For design of a highway sign support structure, an equivalent static load range that produces a static response similar to the dynamic response is used. 2

34 These static load ranges are referred to as the fatigue limit-state load ranges (Dexter, Kaczinski, and Van Dien 1996). The stress ranges calculated using these limit-state design loads are expected to occur at a frequency of 1 in 1, loading cycles, and the structure should be designed so that the stress ranges are below the CAFL for each detail. In the LRFD bridge specifications (AASHTO [1996]), a fatigue design check for large numbers of cycles is implemented based on the infinite-life design approach. The recommended fatigue design loadings in that specification are intended to be representative of the effective load range for the variable amplitude load range histogram. The shape of the truck load-range histograms are such that the limit-state fatigue load range is approximately twice the recommended effective load ranges. Therefore, the fatigue life is considered infinite if the effective stress range calculated using the fatigue design loading is less than half of the CAFL. 21

35 Table 2.1 AASHTO (21) Detail Category A-Values Parameter ET Detail E Detail A 8.1E+7 3.9E+8 CAFL 1.2 ksi 2.6 ksi 22

36 Figure 2.1. S-N Curves for Detail Categories A through ET ( from AWS 2) 23

37 Category E Category ET Figure 2.2. Example of Detail Category E and ET (AASHTO 21) 24

38 CHAPTER 3. STRUCTURAL ANALYSES AND RESULTS 3.1 INTRODUCTION Structural analyses of both VMS structures were performed using a commercially available computer program, GTSTRUDL (22). The finite element capabilities in GTSTRUDL provide structural analysis tools for a wide variety of twoand three-dimensional problems. Static and dynamic analyses were performed to determine the locations of maximum stresses, natural frequencies and modes of vibration of the sign structures. The fatigue design loads defined in AASHTO (21) were used for the static analysis. The static structural analysis results were used in planning the locations of the strain gage instrumentation for the field tests. To verify the accuracy of the structural analyses, an analysis of the structural response due to a known vertical applied loading at midspan was performed on the Birmingham structure. Stresses at the strain gage locations and the displacements of the two bottom chords at midspan were obtained for comparison with the field measurements. The field tests are described in Chapter SPACE FRAME MODELS The VMS support structures consists of two vertical posts at each end and a four-chord box-type truss spanning between the supports. Anchor bolts are used to fix the bottom of the posts to the concrete foundations and U-bolts are used to attach the truss to the WT-section that spans between the vertical posts. Tables 3.1 and 3.2 show the member properties for the Birmingham and Mobile structures, respectively. 25

39 The four chords and posts are API-5L-X52 steel pipes with minimum yield stress of 52 ksi. Struts and diagonals are ASTM A53 with minimum yield stress of 35 ksi. The weights of the VMS panels are 45 lbs and 34 lbs for Birmingham and Mobile, respectively. In GTSTRUDL, three-dimensional beam elements were used to model the cylindrical pipe members of the support structures. Isometric and front views of the three-dimensional models are shown in Figures 3.1 and 3.2. There was an offset between the centroid of the chords and the centroid of the WT-section, so the connection of the truss to the vertical supports was modeled using rigid beam elements to connect these members. The anchor bolts, foundation, and soil conditions were not considered in the analyses. Instead, the base of the vertical supports was defined as the bottom of the post and modeled as completely fixed (i.e., no displacement or rotation was allowed in any of the three principal directions). Analyses for three different static loading conditions were performed first to obtain the stresses in all members. The natural frequencies of the structures were determined for comparison with the free vibration response measured in the field. 3.3 STATIC ANALYSES AASHTO (21) requires that sign support structures shall be designed to resist each of the applicable limit-state equivalent static wind loads acting separately, and the calculated nominal stress range shall not exceed the constant-amplitude fatigue limit (CAFL) values given for a particular connection detail. AASHTO (21) has identified galloping, vortex shedding, natural wind gusts, and truck-induced gusts 26

40 as wind loading mechanisms that can induce large amplitude vibrations and fatigue damage in sign support structures. Dexter and Ricker (22) recommend using only natural wind and truck-induced wind gusts for bridge type support structures. Therefore, only the static loading conditions due to natural wind gust and truckinduced gust were considered. Twenty-six strain gage locations and two deflection points were chosen based upon static analyses of the Birmingham VMS structure due to natural wind gust and truck-induced gust. The stress and deflection results from FEM analyses at 26 strain gage locations and 2 deflection points are shown in this chapter. Similar analyses performed on the Mobile structure resulted in the selection of 14 strain gage locations. The comparisons between results from analyses and field measured data are presented in Chapter Pull Down Loading Condition - Birmingham Structure To assess the accuracy of the structural analyses, a pull down test was performed on the Birmingham structure. As shown in Figure 3.3, static analyses with 2-point vertical loading near the center of the structure were performed. Table 3.3 shows all twenty loading cases of the five trials that were performed during the field tests. The vertical deflections of the two bottom chords at the midspan connection are presented in Table 3.4, and the typical deformed shape is shown in Figure 3.4. The stresses at the 26 strain gage points from the structural analyses are shown in Table 3.5 for trial 2. Results for the other four trials are in Appendix B. These tables show that the stresses in the chords are much higher than the stresses in the 27

41 diagonals. Comparisons of the analysis results to field measured data are presented in Chapter Natural Wind Loads Natural wind gust loading was applied perpendicular to the structure on the vertical projections of the areas of all members and VMS panel according to the following equation: P NW = 5.2 C d I F (psf) (3-1) where C d = drag coefficient, and I F = importance factor. This equation above is based on 5 m/s (11.2 mph) yearly mean wind speed. For locations with more detailed meteorological data, a more precise equation is also provided in the AASHTO (21) commentary. It is assumed here that the Birmingham sign is at a location where the yearly mean wind velocity is 5 m/s. Therefore, equation 3-1 was applied without modification. The importance factor of 1. is used in the design of Category I sign support structures for natural wind gusts. Category I structures are defined as critical cantilevered structures installed on major highways (AASHTO (21)). A drag coefficient of 1.7 is used for a VMS panel, as recommended in Note 7 of Table 3-5 of the AASHTO Specifications (21). Table 3.6 summarizes the dimensions and loads applied to the VMS panels of both structures. Figures 3.5 through 3.8 show the static loading applied to the VMS panel and support structure due to natural wind gust. Tables 3.7 and 3.8 show the stresses at each strain gage location due to the equivalent static loading of natural wind gust. These stresses represent stress ranges due to natural wind gust. In Chapter 4, comparisons are made between these stresses, or stress ranges, and measured 28

42 stress ranges. Here it is interesting to compare the stresses of Tables 3.7 and 3.8 to the constant amplitude fatigue limit (CAFL) for each type of member. For the diagonals and struts the CAFL is 12 psi. The stresses at several members of this type in both the Birmingham and Mobile structures exceed the CAFL. This means that fatigue cracking is expected to eventually occur at the location where these members connect to the main chords, based on the structural analysis results. For the main chords, the CAFL is 26 psi. Stresses exceed this CAFL only in the Birmingham structure at the chord splices near midspan Truck-Induced Gusts Loading As stated in the literature review, VMS structures are particularly susceptible to truck-induced wind gusts because of their thickness in the direction parallel to traffic flow that is exposed to this vertical pressure. The fatigue load due to truckinduced gusts provided in AASHTO (21) is: P TG = 18.8 C d I F (psf) (3-2) where C d = drag coefficient, and I F = importance factor. The Birmingham, AL. structure is erected at a location where the posted speed limit is less than 65 mph. However, Equation 3-2 was applied without modification to be conservative. The importance factor (I F ) of 1. is used in the design of Category I sign support structures for truck-induced gusts. The drag coefficient of 1.7 should be used for a VMS panel, as is recommended in Note 7 of Table 3-5 in AASHTO (21). Therefore, a summary of the equivalent static truck-induced gust load to be applied to each VMS panel is presented in Table

43 The equivalent static load to be applied to each member in the bottom plane of the truss directly behind the VMS panel is summarized in Table 3.1. Figures 3.9 through 3.12 show the static loading condition of the VMS panels and support structures due to truck-induced gust. Tables 3.11 and 3.12 show the stresses at the strain gage points due to the equivalent static load of truck-induced gust. At most strain gage locations the stresses due to natural wind gusts were larger than these dues to natural wind gusts were larger than those due to truckinduced gusts. A comparison of the combined axial and bending stresses due to these two types of loadings (Table 3.7 verses Table 3.11 and Table 3.8 verses Table 3.12) shows that the stresses due to truck-induced gusts are the largest at only two locations, V3 and V4 in the Birmingham structure. An inspection of the stresses due to the truck-induced gust shows that the CAFL is exceeded only at strain gage location V3 of the Birmingham structure. There are additional locations on diagonals and struts where the stresses are near the CAFL, above 1 psi. 3.4 DYNAMIC ANALYSES A natural frequency analysis was conducted to establish the dynamic characteristics of the sign support structures and for comparison of the predicted free vibration response with recorded field data. The VMS panels were modeled as a series of framed beam elements and nodes offset from the main chords (see Figures 3.13 through 3.2). The nodes and offset represents the centroid and eccentricity of the VMS panels. The mass of the sign was distributed equally among these nodes. 3

44 3.4.1 Birmingham Structure The first two mode shapes are shown in Figures 3.13 and The first mode involves out-of-plane bending of the truss (i.e., displacement in the horizontal plane). The frequency of the first mode is 1.35 Hz. The second mode involves inplane bending (i.e. displacement in the vertical plane). GTSTRUDL predicted a frequency of 2.2 Hz for the second mode. Figure 3.15 shows the third mode shape of the Birmingham support structure. For the third mode, the sign structure moves side to side where bending of the supports occurs. The frequency is 2.22 Hz. The fourth mode is shown in Figure As shown in this figure, the fourth mode involves twisting of the structure about the axis of the truss, resulting in an out-ofplane motion. The frequency of this mode is 4.74 Hz. A list of the first six modal frequencies is presented in Table Mobile Structure The first two mode shapes are shown in Figures 3.17 and In the first mode, the structure moves side to side, and has a frequency of 1.28 Hz. The second mode of vibration involves out-of-plane bending of truss members (i.e. moving from front to back). The frequency determined for the second mode is 3.1 Hz. The third and fourth modes of vibration are shown in Figures 3.19 and 3.2. As shown in these figures, the third mode involves vibration in the vertical direction (i.e. up and down), and the fourth mode of vibration involves twisting motion of the truss. The frequencies of the third and fourth modes of vibration are 3.1 Hz and 8.5 Hz, respectively. A list of the first five modal frequencies is presented in Table

45 3.5 CONCLUSIONS Three-dimensional models of the VMS structures were created using GTSTRUDL. Natural wind gust and truck-induced wind gust fatigue design loads were applied in accordance to Section 11 of AASHTO (21). The structural analyses predict stress ranges above the CAFL at various locations in both the Birmingham and Mobile structures. This means that these structures are predicted to have finite fatigue lives based on these analyses. A field study was conducted (see Chapter 4) to investigate the response of the structures to natural conditions. Field tests were performed to assess the accuracy of the structural analyses and monitor stress ranges over a period of time. A finite life check and fatigue life evaluation were also performed using the measured data. The results of the field study are presented in Chapter 4 with comparisons to the analytical results of this chapter. 32

46 Table 3.1 Member Properties Birmingham Structure Member Type Outside Diameter (OD, inch) Thickness (t p, inch) Length (L, inch) Chord Truss Support Strut Diagonal Post Vertical Plane Horizontal Plane Vertical Plane Horizontal Plane Plane Perpendicular to Axis of Truss Left Right Diagonal Table 3.2 Member Properties Mobile Structure Member Type Outside Diameter (OD, inch) Thickness (t p, inch) Length (L, inch) Chord Truss Strut Vertical Plane Horizontal Plane Vertical Plane Support Diagonal Post Horizontal Plane Plane Perpendicular to Axis of Truss Left Right Strut Diagonal

47 Table 3.3 Pull Down Loading Cases Trial Load (pounds) Trial Load (pounds)

48 Table 3.4 Deflection due to Pull Down Loading Cases Trial Applied Load (pound) Deflection (inch) Trial Applied Load (pound) Deflection (inch) Front Bottom Chord Back Bottom Chord

49 Member Type Diagonal & Strut Front Top Chord Back Top Chord Back Bottom Chord Gage Name Table 3.5 Structural Analysis Results for Trial 2 23 lbs 31 lbs 22 lbs Axial Bending Total Axial Bending Total Axial Bending Total S S V V H H H H H H V V WT T Section T

50 Table 3.6 Natural Wind Gust Loads Applied to VMS Panels VMS Panel P NW (psf) (5.2x C d x I F ) Height (H) (ft) Width (W) (ft) F NW (lbs) (P NW x H x W) Birmingham Mobile Table 3.7 Stresses at Strain Gage Locations due to Natural Wind Gust Birmingham Structure Member Type Gage Name Axial Bending Total S S V V H Diagonal & Strut H H H H H V V Front Top Chord Back Top Chord Back Bottom Chord WT T Section T

51 Member Type Table 3.8 Stresses of Strain Gage Points Due to Natural Wind Gust Mobile Structure Gage Name Axial Bending Total H H Diagonal & Strut H H V V Front Top Chord Back Bottom Chord Table 3.9 Truck-Induced Gust Loadings Applied to VMS Panels VMS Panel P TG (psf) (18.8 x C d x I F ) Depth (D) (ft) Width (W) (ft) F TG (lbs) (P TG x H x W) Birmingham Mobile Table 3.1 Truck-Induced Loadings Applied to Truss Members Behind VMS Structure Birmingham Mobile Member P TG (psf) (18.8 x C d x I F ) Outside Dia. (OD) (ft) Length (L) (ft) F TG (lbs) (P TG x OD x L) Chords Diagonals Chords Diagonals

52 Table 3.11 Stresses at Strain Gage Locations due to Truck-Induced Gust Birmingham Structure Member Type Diagonal & Strut Front Top Chord Back Top Chord Back Bottom Chord Gage Name Axial Bending Total S S V V H H H H H H V V WT T Section T

53 Table 3.12 Stresses of Strain Gage Points Due to Truck-Induced Gust Mobile Structure Member Type Diagonal & Strut Top Front Chord Bottom Back Chord Gage Name Axial Bending Total H H H H V V Table 3.13 Modal Frequencies Birmingham Structure Mode Natural Circular Frequency (Rad/Sec) Natural Frequency (Cyc/Sec, Hz) Natural Period (Sec/Cyc)

54 Table 3.14 Modal Frequencies Mobile Structure Mode Natural Circular Frequency (Rad/Sec) Natural Frequency (Cyc/Sec, Hz) Natural Period (Sec/Cyc)

55 145 ft Sign Panel (a) Isometric View 145 ft 33 ft in 39 ft in 72 ft - 6 in 3@1.68ft + 15in 4@9.48ft + 15in 5 ft 15 ft - 1 in. Sign Panel = 3.75 ft (b) Front View Figure 3.1. Model of VMS Support Structure on US-28 in Birmingham 42

56 89 ft - 3 in. Sign Panel (a) Isometric View 89 ft - 3 in. 29 ft - 9 in. 29 ft - 9 in. 29 ft - 9 in. 3 ft - 6 in. 24 ft in. Sign Panel = 24 ft 22 ft in. (b) Front View Figure 3.2. Model of VMS Support Structure on I-1 in Mobile 43

57 Back o o Front (a) Isometric View o (b) Front View Figure 3.3. Pull Down Loading Condition of Birmingham Support Structure 44

58 Measured Deflection Point (a) Isometric View Measured Deflection Point (b) Front View Figure 3.4. Typical Deformed Shape of Birmingham Structure Due to Pull Down Loading Condition 45

59 36 in. W6x9 23 in lbs 231 lbs VMS 54 in. VMS 12 in in lbs Figure 3.5. Static Loading Condition of Birmingham VMS Due to Natural Wind Gust o o Figure 3.6. Static Loading Condition of Birmingham Support Structure Due to Natural Wind Gust 46

60 19 in. W6x9 32 in lbs VMS 42 in. VMS 977 lbs 11 in. 32 in. 977 lbs Figure 3.7. Static Loading Condition of Mobile VMS Due to Natural Wind Gust o o o o o o o o o o o o o o o o ooo o o o o o o ooo o o o o o o o o o o o o o o o o o ooo o o o o o o ooo o o o o o o o o o o o o o o o oo o o o o o o o Figure 3.8. Static Loading Condition of Mobile Support Structure Due to Natural Wind Gust 47

61 36 in. W6x9 23 in. 626 lbs 12 in. VMS 54 in. VMS 21.5 in. 626 lbs 1151 lbs 32.2 in lbs 1151 lbs Figure 3.9. Static Loading Condition of Birmingham VMS Due to Truck-Induced Gust o o o o o o Figure 3.1. Static Loading Condition of Birmingham Support Structure Due to Truck- Induced Gust 48

62 19 in. W6x9 32 in. 11 in. VMS 42 in. VMS 388 lbs 32 in. 388 lbs 67 lbs 26.8 in. 67 lbs 67 lbs Figure Static Loading Condition of Mobile VMS Due to Truck-Induced Gust o o o o o o o o o Figure Static Loading Condition of Mobile Support Structure Due to Truck-Induced Gust 49

63 Back Front Figure First Mode Shape of Birmingham Support Structure Top Bottom Figure Second Mode Shape of Birmingham Support Structure 5

64 Side Side Figure Third Mode Shape of Birmingham Support Structure Figure Fourth Mode Shape of Birmingham Support Structure 51

65 Side Side Figure First Mode Shape of Mobile Support Structure Back Front Figure Second Mode Shape of Mobile Support Structure 52

66 Top Bottom Figure Third Mode Shape of Mobile Support Structure Figure 3.2. Fourth Mode Shape of Mobile Support Structure 53

67 CHAPTER 4. FIELD TESTS AND FATIGUE EVALUATION 4.1 INTRODUCTION The primary goal of the field tests was to measure nominal stress ranges in representative members of the sign support structures due to wind gusts to determine if these members were susceptible to fatigue cracking in the future. A static pull down load test was performed on the Birmingham structure to investigate the accuracy of a standard space frame structural analysis of the tubular superstructure by allowing direct comparisons of the measured and calculated stresses and displacements. Computer analyses using GTSTRUDL were performed to identify the high stress locations due to design fatigue loads specified in the AASHTO Standard Specifications for Structural Supports for Highway Signs, Luminaires and Traffic Signals (AASHTO 21). Strain gages for making field measurements were installed in these high stress locations. An anemometer was utilized to measure wind speed and direction during the testing period. Work began in late March 23 to install the strain gages, wind anemometer, and data acquisition system on the VMS structure in Birmingham. Long-term monitoring began on April 24 th and was suspended from April 29 th through May 12 th while a catwalk was installed on the structure by others for a purpose unrelated to this research. Monitoring resumed on May 13 th and continued through June 11 th. There were approximately three days during this time when the system shut down due to insufficient power supply. 54

68 Work began in mid-september 23 to install the instrumentation on the Mobile structure. Long-term monitoring started on September 18 th and ended on January 1 th, 24. Strain data was collected for approximately 12 weeks during this period. There were a few weeks during this period where the data acquisition program malfunctioned, and this data was omitted from further analysis. 4.2 INSTRUMENTATION AND DATA ACQUISITION Strain Gage Locations Birmingham Structure The support structure is geometrically similar on each side of the chord splices at midspan. Strain gages were installed only on the east/right half. Figure 4.1 illustrates the locations of the strain gages. A total of 26 electrical resistance strain gages were installed on 1 members of the structure. These members provide a representative sample of locations where fatigue cracking may occur and include three of the four main truss chords, diagonals and struts in the truss and supports, and the horizontal WT-section on which the truss is supported. Once these were identified by structural analysis as highly stressed members, a closer review of the analysis results was made to identify where the largest stresses occurred (i.e. top, bottom, front, or back sides and at which end of the member). Gage locations on the structure were chosen to measure nominal stress ranges for use in predicting fatigue life. According to AASHTO (21), the nominal stress ranges are not to exceed the constant amplitude fatigue limit so that infinite fatigue life is achieved. Since the nominal stresses are highest at member ends, the 55

69 gages were located as close to the ends as possible. However, stress concentrations are prevalent near the welds at the connections; therefore, the gages were installed a short distance away from the member ends as illustrated in Figure 4.2 to avoid the stress concentrations. On the diagonal and strut numbers, the strain gages were placed two outside member diameters from the member ends. Two gages were installed on each of the diagonal and strut members on diametrically opposite sides of the circular cross sections (i.e. top and bottom and/or front and back sides of the member). Four gages were installed on each of the main chords at intervals 9 degrees apart around the circumference of the cross section. These strain gages were installed 4 in. from the toe of the welds around the member at the chord splices. Two gages were installed on the WT-section as shown in Figures 4.3. Due to the location of the truss chords bearing on the WT-section (see Figure 1.7), these gages were placed 1 inch from the end of the member, too close to avoid the stress concentration effect Mobile Structure The Mobile structure had a shorter span length compared to the Birmingham structure, and all members were accessible for instrumentation. A total of 14 strain gages were installed on 5 members on the structure as shown in Figure 4.4. Two of the four chords at a splice connection, and three diagonal members were instrumented in the same manner as the Birmingham structure Strain Gages Only one type of strain gage was used on the two structures. The gage type was an EA series bondable strain gage from Measurements Group, Inc. with an 56

70 electrical resistance of 35 ohms. The strain gage installation process consisted of removing the galvanizing with an electric grinder to expose the steel. The steel was then sanded with fine grain sandpaper until the desired smoothness was reached. The exact gage location was then marked using a ballpoint pen. The steel surface was then cleaned with a chemical solvent to remove all oil and debris that would prevent the gage from completely bonding with the steel. The gage was then bonded to the surface using an M-Bond adhesive from Measurements Group, Inc. After the installation process, the electrical resistance of the gage was verified to assure no damage to the gage occurred during installation. Shielded lead wires were then connected from the gages to the data acquisition system located at the right support Wind Anemometer A wind anemometer (model 513V) by RM Young Company was used to continuously record wind speed and direction. This device was monitored with the data acquisition system so wind data could be stored with the strain data in the datalogger. The anemometer was programmed to output wind speed and direction in miles per hour and degrees using a three second averaging time. A record was produced every three seconds indicating the average wind speed and direction over the previous three second period. The anemometer was powered by an external 12 volt battery Data Acquisition System A CR9 data logger from Campbell Scientific, Inc. measured and recorded the stress ranges at all strain gages on the structure. Two external 12 volt batteries 57

71 powered the data logger. The batteries allowed the system to run unattended for approximately six days; therefore fully charged batteries were swapped out every six days. Recorded data was downloaded from the logger s internal memory card each time the batteries were swapped out to ensure the program was running properly. The data acquisition system recorded data according to the program instructions input by the operator. The type of data recorded varied depending on the test performed. Time history data was recorded during the calibration and random truck tests and during the pull down test. During continuous monitoring, the system was programmed to perform a rainflow counting algorithm that recorded strain cycles and output a strain cycle histogram every 15 minutes for each gage. The logger stored the individual strain cycle magnitudes in histograms with interval sizes of 8 microstrain. For all strain gage measurements, the data logger was set at a sampling rate of 83.3 Hz. 4.3 FIELD TESTS Several controlled tests were performed after the instrumentation was in place to verify the performance of the strain gages and wind anemometer. Time history strain data was recorded in short time increments on several occasions while the structure was stationary to determine the range of electronic noise in the gages. The noise in the gages was determined to be approximately 5 microstrain, and measured cycles of this magnitude were not counted in the rainflow counting algorithm. The controlled tests revealed that gages H2 and Back Bottom Chord No. 2 (see Figure 4.1) on the Birmingham structure were not operating properly, but due to 58

72 time constraints these gages were not replaced. Data from these gages are not reported here. Controlled tests of the Mobile structure and data acquisition system revealed that gage FT3 (see Figure 4.4) was not operating properly, but the gage was not replaced due to time constraints. Data from that gage are not reported here Response to Truck-Induced Gusts Birmingham Structure By observation, large trucks passing beneath the VMS produce gusts causing vibration of the support structure. It was not feasible to perform an extensive series of controlled tests to study truck-induced gust loadings. But, time history data was recorded for three passes of large trucks. For each pass, the truck was traveling in a different lane at an estimated speed of close to the posted speed limit of 45 mph. Time history data recording began just before the truck passed beneath the sign and ended a few seconds after the truck had passed. Based on the recorded data and observations throughout the course of this project, the severity of the vibration appears dependent on truck speed, lane in which the truck was traveling, and type of truck. The VMS is centered over all three northbound lanes of traffic, but it only entirely covers the center lane. Therefore, trucks traveling in the middle lane produce more vibration of the structure than trucks traveling in the outer lanes. Figures 4.5 through 4.7 show the time history response to a truck traveling in the middle lane for gages H1, FT3, and T1. This data provides comparisons with the structural analysis of the dynamic characteristics of the structure. A natural frequency of 1.35 Hz in the 59

73 first mode of vibration is observed from the data which is consistent with the analysis results presented in Chapter 3. Truck speeds at this location are typically lower than the posted speed limit because of the congested traffic conditions common at this location and the presence of traffic signals on either side of the structure. These traffic conditions prevent trucks from traveling underneath the sign at high speeds that would produce large wind gusts during most of the day. The recorded data shown in Figure 4.5 through 4.7 is believed to be representative of a relatively large truck-induced loading for this structure. The truck-induced strain data indicated the chord members and diagonal members experienced stress ranges of approximately.87 ksi and.58 ksi, respectively, which are 33 percent and 48 percent of the CAFL for the chord and diagonal members, respectively. Table 4.1 shows comparisons between the maximum measured stress ranges due to a single truck and the stress ranges calculated using the truck-induced design load from AASHTO (21). The stress ranges measured at the chord splices are less than the calculated values. The stress ranges measured at some of the diagonals and struts are larger than the calculated stress ranges. As mentioned in Chapter 3, AASHTO (21) specifies an equivalent static load condition to represent truck-induced wind gusts. According to AASHTO (21), the truck-induced loads are applied vertically to the bottom side of all structural members and attachments resulting in a vertical deflection of the structure. Therefore the stresses obtained from the analysis are representative of these vertical loading and deflection conditions. However, the observed deflections in the field due 6

74 to truck-induced gusts are primarily in the horizontal direction. This is apparent because the first mode of vibration for the structure is dominated by horizontal motion (see Figure 3.13). Not including a horizontal component of the truck-induced gust loading in the analysis is probably the reason that the measured stress ranges exceed the calculated values. However, this is not a point of concern with AASHTO (21) because a structure must also be designed for significant horizontal loading due to natural wind gusts Mobile Structure Large trucks passing beneath the sign at high rates of speeds did not cause visible vibrations of the structure, but some truck-induced gust data was recorded. Time history data was recorded for three conditions of trucks passing underneath the VMS: a single truck; 2 trucks traveling side-by-side, and 3 trucks in a row traveling in the same lane. For all conditions, the trucks were estimated to be traveling between 7 and 75 mph. Strain recorded for all three conditions revealed the stress ranges were very low due to the truck-induced gusts. Figure 4.8 shows the response of gage V1 to the single truck condition. This particular gage had the highest predicted stress range from the structural analysis (1.1 ksi); however, the largest measured stress range due to the passing of a single truck for this member was approximately.33 ksi. The measured truck-induced data indicated none of the gages measured stress ranges above the CAFL. Table 4.2 shows comparisons between the maximum measured stress ranges due to a single truck and the stress ranges calculated using the truck-induced gust design load from AASHTO (21). The 61

75 measured stress ranges at gages on some of the diagonal members are above the calculated values by a small amount. The stress ranges measured in the chords are below the calculated values. The truck-induced strain data also provided comparisons with the structural analysis of the dynamic characteristics of the structure. A natural frequency of 3. Hz was determined from the free vibration response. This is in agreement with the natural frequency of the second mode of vibration obtained from the analysis results in Chapter 3. Since the first mode was characterized by side-to-side motion, the strain gages could not accurately measure the response in this direction Static Pull Down Test Birmingham Structure The objective of the static pull down loading test was to investigate the accuracy of a standard space frame structural analysis of the tubular structure. To achieve this, load increments were applied to the structure and measurements of the resulting strains and displacements were made. Prior to the static tests, structural analyses were performed to verify that the structure would not be stressed beyond design allowable stresses by a test load of 3 lbs. The structural analysis indicated that the structure had more than adequate strength to support the test loading. Maximum stresses due to the test loading were less than 8% of the yield stress and less than 13% of the allowable stress. In most members the stresses were significantly less. In preparation for this test, a load cell was fabricated, instrumented with strain gages, and calibrated in the lab using a load testing machine and a strain indicator. A linear relationship between measured strain and applied load was established 62

76 during calibration. The load cell was then implemented in the field test to measure the applied loads. The deflections at the midspan connection were measured using a deflectometer attached from the ground to each bottom chord using a small diameter steel wire. Strain gages were also installed on the deflectometers, and a relationship between measured strain and deflection was established. Figures 4.9 through 4.12 are illustrations of the test setup Testing Procedure The static load test consisted of applying a vertical load near midspan of the structure by using a cable, a pulley, and a hand winch attached to a stationary truck parked in the median. During this test, one person manually loaded and unloaded the structure using the hand winch while a second person recorded the applied load from the strain reading of the load cell. There was also an operator on the shoulder of the highway recording strain data for all of the gages using a laptop computer connected to the data logger. The data logger was programmed to record time history data throughout the entire duration of the test. For all strain measurements, the data logger sampled the strain gages continuously at 2 Hz and output an average reading at 2 Hz. After each load increment, time history data was recorded while holding the load constant. Therefore there were separate data files for each load along with a data file containing all the data for the entire test. A total of five trials of the static pull down test were performed. The structure was loaded incrementally in each trial, and data was recorded at the end of each increment. Each trial had slightly different load increments due to the lack of precision of the hand winch. 63

77 For each trial, initial zero load data was recorded before loading began. The load was then increased by jacking down on the hand winch. At the end of each load increment, the operator was signaled to record data while the load was held constant. Data was recorded for approximately 45 to 6 seconds after each load increment. Once the target maximum load was reached, the loading was decreased and held at a minimal load before the structure was completely unloaded. For the first trial, there were ten loading increments. Due to a lack of battery life of the laptop computer that controlled the data acquisition system and concerns about the potential for errors due to thermal effects, the remaining four trials consisted of 5 increments each. Trial 1 required approximately 15 minutes. The total loading and unloading time of trials 2 through 5 was reduced to approximately 4 minutes each. Table 4.3 shows the loading increments for each trial. Figure 4.13 is a graphical representation of two strain gage outputs for the entire duration of trials 2 through Data Reduction As explained in the testing procedure, the trials consisted of recording data at each load increment for a period of approximately 45 to 6 seconds. At an output frequency of 2Hz, between 9 and 12 strain readings were output at the end of each load increment including the zero load increment at the beginning and end of each trial. The data reduction process involved taking an average strain reading for each gage over the time period between load increments and adjusting these values by subtracting the initial average strains at zero load. This was necessary to eliminate 64

78 the offset in the initial gage readings at the beginning of the trial. This adjustment was performed for each trial. Figure 4.14 illustrates this process Results Figures 4.15 through 4.39 graphically compare the predicted and measured stresses and deflections for trials 1 and 2 of the static pull down loading test. Plots for trials 3 through 5 are presented in the Appendix C. Overall, the stresses and deflections measured in trials 2, 3 and 4 match best with the structural analysis results. Plots for trial 1 in Figures 4.15 through 4.39 show some nonlinear behavior during this trial. The nonlinear responses are believed to have resulted from seating of the bolted connections and limited local yielding at the welded connections of the tubular members. Conditioning of the structure by applying one cycle of the test loading prior to recording data would have eliminated the confusing data, but this did not appear to be the safest approach to performing the tests. The stresses in the chords were the largest measured responses and matched very well with the stresses calculated from by the structural analyses. Table 4.4 provides a comparison of measured and calculated stresses in the chords and vertical deflections at midspan for the highest load applied in each trial. The values listed indicate that the calculated values were typically larger than the measured values. Ratios of the measured stress to the calculated stress are shown for each measurement. Percentage differences range from the measured stresses being 7% above the calculated stress to the measured stress being 5% below the calculated value. For all the chord stress measurements the average percentage difference 65

79 was 1% with 85% of all of the measurements being with 25% of the calculated values. The plots of Figure 4.26 through 4.39 show significant deviations from the predicted stresses in members that experienced low levels of stress. This is apparent in many of the diagonal and strut members where the maximum measured stresses were typically below 2 psi. Strain gage locations V1, H1, H5 and T1 experienced stresses of opposite sign to those predicted Long-Term Monitoring Accurate calculation of the effective stress range in each member required large samples of stress ranges from normal daily loading conditions. During the longterm monitoring, the CR9 datalogger continuously recorded strain ranges occurring from normal daily truck traffic and natural winds using a rainflow cycle counting algorithm. This provided a record of the number and magnitude of all significant strain cycles that occurred during the monitored period. The stress ranges measured during long-term monitoring represented a combination of truck-induced and natural winds. The number and magnitude of stress cycles counted could not be separated into truck-induced stress ranges and natural wind stress ranges. However, the results of the random truck test indicate that stress ranges experienced in all instrumented members were well below the CAFL. Higher stress ranges were measured when periods of relatively high wind speeds were recorded by the wind anemometer. It is believed that the majority of the stress cycles counted during the long-term monitoring is a result of natural winds. 66

80 Birmingham Structure Strain gage and anemometer data was collected for a total of 31 days. During this period, stress ranges above the CAFL were recorded in several members. Table 4.5 lists for each strain gage the total number of cycles recorded (above the threshold of 2% of the CAFL) and the percentage of those cycles that were above the CAFL. A CAFL of 1.2 ksi is shown in Table 4.5 for both strain gage locations on the WT section in the vertical support. These two strain gages were located 1 inch from the weld at the end of the member, and the measurements from these gages include a stress concentration effect. Hence, the CAFL listed, 1.2 ksi, is a lower bound for comparison with the measurements. Generally, stress ranges above the CAFL were recorded only at times when the natural wind speed was above 15 mph. The only exception was for strain gage S1 on the diagonal member of the vertical support. Stress ranges above the CAFL were measured at this gage when the wind speed was below 15 mph. The recorded wind data indicates that the prevailing wind direction is out of the north-northwest direction. This prevailing wind direction is approximately perpendicular to the VMS panel and creates a worst case wind loading condition. Figure 4.4 shows a period of strong winds during a thunderstorm when high stress cycles were recorded. Each data point in Figure 4.4 represents a 3 second average wind speed Mobile Structure Data was collected for a period of approximately 82 days. Very few cycles above the CAFL were recorded during this period even though wind gusts exceeding 67

81 2 mph were recorded on several occasions. A summary of the measurements is provided in Table 4.6. The wind anemometer data revealed that the wind direction was variable on any given day which is expected for locations near the coast Stress Range Occurring at 1 in 1, Cycles As previously mentioned, an infinite life fatigue design approach is currently used in the design of highway sign, signal, and luminaire support structures per AASHTO (21). The normal loading conditions on the sign structure due to truck traffic and natural winds produce a variable amplitude history of stress ranges. As long as the highest stress ranges in the history occur at a frequency of less than 1 in 1,, the sign structure should not experience fatigue cracking, but should have infinite life. Hence it is necessary to estimate the stress range that was measured at a frequency of 1 in 1, for use in performing a finite life check. This stress range is also referred to as the limit state stress range. Tables 4.5 and 4.6 show the results for each structure including the total number of cycles counted with magnitude above 2 percent of the CAFL and the maximum stress range that occurred during the monitoring period. Birmingham Structure. Table 4.7 provides a side by side comparison of the calculated and measured limit-state stress ranges for the Birmingham structure. The stress range values represent total stress (i.e. axial and bending stress). The results show that the majority of the measured limit-state stress ranges were below the calculated values, but not all. Four of the gages did not measure more than 1, loading cycles during the monitoring period, therefore the stress range occurring at a 1 in 1, frequency could not be accurately determined from the recorded data. 68

82 Mobile Structure. Table 4.8 presents the comparisons of the calculated and measured limit-state stress ranges for the Mobile structure. During the monitoring period, less than 1, loading cycles were recorded in all gages on the two main chords. Thus, the stress range occurring at a frequency of 1 in 1, loading cycles could not be accurately determined from the recorded data. As shown in this table, all measured limit-state stress ranges in the diagonal members were less than the calculated values Effective Stress Range A stress range histogram was required for the calculation of the effective stress range of each strain gage. The numbers of stress cycles measured at each gage during the long-term monitoring were extrapolated to estimate the number of cycles expected per year. These numbers of cycles per year are listed in Tables 4.9 and 4.1. Only stress cycles at or above 25 to 3 percent of the CAFL contribute to fatigue crack initiation and growth. Therefore measured stress cycles of small amplitude were omitted from the effective stress range calculations. The cutoff point used in the calculations was 27 percent and 2 percent of the CAFL for the truss chord splices (category E ) and diagonals (category ET), respectively. Tables 4.11 and 4.12 list the effective stress ranges from the field measurements at each strain gage. These effective stress ranges converged to a reasonably constant value after only a few days of measurements at each location. To illustrate the variability of the measurements, effective stress ranges calculated for individual days during the monitoring period are plotted in Figures 4.41 through The plots show that even the effective stress ranges for each day were reasonably 69

83 constant. This provides confidence that the effective stress ranges calculated for the entire monitoring period are representative of the stress ranges experienced by the sign structures. Table 4.11 lists for each strain gage on the Birmingham structure the measured effective stress range, limit-state stress range and the maximum stress range measured during the entire monitoring period. This table also includes ratios of limit-state to effective stress range and maximum stress range to limit state stress range. The ratio of the measured limit-state stress range to measured effective stress range was reasonably consistent for all gage locations and had an average of 3. The ratio of maximum stress range to limit state stress range varied from 1. to 2.8. Table 4.12 lists for each strain gage on the Mobile structure the measured effective stress range, limit-state stress range and the maximum stress range measured during the entire monitoring period. The ratio of measured limit-state stress range to measured effective had an average of 2.2. The ratio of maximum stress range to limit state stress range varied from 1. to FATIGUE EVALUATION Finite Life Check A finite life check was performed by comparing the measured stress ranges occurring at a frequency of 1 in 1, loading cycles to the respective CAFL for each member. The required data are shown in Tables 4.5 and 4.6. Table 4.5 shows that five of the diagonal and strut members on the Birmingham structure did 7

84 experience limit-state stress ranges above the CAFL. Therefore, the structure has finite life and is susceptible to fatigue cracking. Table 4.5 for the Mobile Structure shows that no limit-state stress ranges above the CAFL were measured during the monitored period. Hence, the Mobile structure is predicted to have infinite life. A fatigue life evaluation was conducted on both structures to estimate when fatigue cracking may be expected to occur. The fatigue life evaluation was made for the Mobile structure for completeness even though that structure is predicted to have a practically infinite life. The following sections discuss the process of using recorded data to perform the fatigue life evaluation Fatigue Life Evaluation Fatigue lives were calculated using the effective stress ranges measured at each strain gage location in each structure. These fatigue lives were calculated using the detail constants and procedures outlined in the literature review of Chapter Two. The number of stress cycles measured at each gage during the long-term monitoring were extrapolated to estimate the number of cycles expected per year. This allowed the fatigue lives to be reported as a number of years instead of a number of cycles. The results for the Birmingham and Mobile structures are listed in Table 4.13 and 4.14, respectively. As stated in the previous section, finite fatigue life is expected at only five of the strain gages locations, all of which are in the Birmingham structure. Results for these five strain gage locations are shown in bold print in Table Fatigue lives at the other strain gage locations are reported to illustrate what fatigue live is expected if it is assumed that the finite life check was in error. This was done because the length of long-term monitoring performed here is 71

85 believed to have provided more accurate measurements of the effective stress range than the stress ranges occurring at a frequency of 1 in 1,. The fatigue lives shown in Tables 4.13 and 4.14 are the total fatigue life from the time these structures were new. Since both structures are relatively new, these total fatigue lives are essentially equal to the remaining life. As indicated in Tables 4.13 and 4.14, the support diagonal (gage S1) in the Birmingham structure is the only member in which cracking is predicted prior to the 5-year design life of the structure. There are three other locations in the Birmingham structure and two in the Mobile structure where the calculated fatigue lives are between 5 and 6 years. 4.5 CONCLUSIONS Based on the structural analyses of the previous chapter, several members of the Birmingham and Mobile structures were predicted to have finite fatigue life. The primary objective of the field tests was to collect data for an alternate fatigue evaluation of these structures. A static load test was performed to investigate the accuracy of a standard space frame structural analysis of the tubular sign structure. Stresses and deflections due to known applied loads were measured and compared to calculated values. Overall the structural analysis results and measured values agree more closely than is common for highway bridges. Stresses at the highest applied load level in the main truss chords near the midspan splice were generally between 2 psi and 3 psi. Stresses calculated in the structural analysis were typically larger 72

86 than the measured values. Percentage differences ranged from the measured stresses being 7% above the calculated stress to the measured stress being 5% below the calculated value. For all the chord stress measurements the average percentage difference was 1% with 85% of all of the measurements being with 25% of the calculated values. The stresses in the strut and diagonal members were generally below 2 psi. The stresses measured in some of these members matched very well with the calculated stresses, and others did not. At four strain gage locations the measured stresses were opposite sign from the calculated values. The differences between the measured and calculated values for the struts and diagonals is primarily the result of the low magnitude of stresses in these members. Overall, the results from a standard structural analysis of a sign structure such as the one tested should be sufficiently accurate for fatigue design of the structure. The most significant source of conservatism, or lack of conservatism, in an analytical fatigue design or evaluation of this type structure is in defining the applied wind loads. The specific goal of monitoring each sign structure for an extended period of time was to measure stress ranges occurring under normal wind and traffic conditions and estimate yearly stress cycle histograms. Long-term monitoring of the applied stress cycles indicated that several members in the Birmingham VMS structure experienced limit-state stress ranges (occurring at 1 in 1, cycles) above the CAFL, therefore this structure has finite life and is susceptible to fatigue cracking. No members in the Mobile structure experienced stress cycles above the CAFL. The stress cycle histograms were used in the fatigue life evaluations to 73

87 determine the fatigue life at each of the strain gages locations. These fatigue life estimates indicate that only one member, a diagonal in the support of the Birmingham structure, is predicted to have a fatigue life of less than the 5-year design life. 74

88 Table 4.1 Comparisons of ulated and Measured Truck-Induced Stress Ranges Birmingham Structure Member Type Diagonals & Struts Front Top Chord Back Top Chord Back Bottom Chord WT-Section Gage S r (ksi) ulated Measured S S V V H H H H H V V T T Table 4.2 Comparisons of ulated and Measured Truck-Induced Stress Ranges Mobile Structure Member Type Diagonals Front Top Chord Back Bottom Chord Gage S r (ksi) ulated Measured V V H H H H FT FT FT BB BB2.6.3 BB BB

89 Table 4.3 Static Pull Down Load Increments for Each Trial Static Field Test (6/11/3) Trial 1 Trial 2 Trial 3 Trial 4 Trial 5 Loading Increment Total Loading Increment Load (lb) Loading Increment Load (lb) Loading Increment Load (lb) Loading Increment Load (lb)

90 Table 4.4 Peak Load Comparisons at Main Chord Members (a) Stresses (psi) Trail 1 Trial 2 Trial 3 Trail 4 Trail 5 Member Gage 269 lbs Meas. Meas./.. Meas. Meas./.. Meas. Meas./.. Meas. Meas./.. Meas. Meas./ Front Top Chord Back Top Chord Back Bottom Chord (b) Deflections (in.) Member Front Chord Back Chord Trail 1 Trial 2 Trial 3 Trail 4 Trail lbs Meas. Meas./.. Meas. Meas./.. Meas. Meas./.. Meas. Meas./.. Meas. Meas./

91 Table 4.5 Field Measured Stress Cycle Summary Birmingham Structure Member Type Diagonals & Struts Front Top Chord Back Top Chord Back Bottom Chord WT Section Gage CAFL (ksi) Total Number of Cycles Above 2% of CAFL % of Cycles Above CAFL S r (ksi) Occurring at 1 in 1, Cycles S r max (ksi) S S V V H H H H * 3.36 H V V * * * T1 1.2** T2 1.2** * Not available **Lower bound for comparison to measured stress ranges 78

92 Table 4.6 Field Measured Stress Cycle Summary Mobile Structure Member Type Gage CAFL (ksi) Diagonals Front Top Chord Back Bottom Chord Total Number of Cycles Above 2% of CAFL % of Cycles Above CAFL S r (ksi) Occurring at 1 in 1, Cycles S r max (ksi) V V H H H H FT FT FT BB BB BB BB

93 Table 4.7 Comparisons of ulated and Measured Limit-State Stress Ranges Birmingham Structure Member Type Diagonals & Struts Front Top Chord Back Top Chord Back Bottom Chord WT-Section S r (ksi) Gage (1 in 1, Cycles) S r,calc /S r,meas ulated Measured S S V V H H H H5.81 ** ** H V3 1.36* V4 1.11* ** ** ** ** ** ** T T * Truck-induced ** Not Available 8

94 Table 4.8 Comparisons of ulated and Measured Limit-State Stress Ranges Mobile Structure Member Type Diagonals Front Top Chord Back Bottom Chord S r (ksi) Gage (1 in 1, Cycles) S r,calc /S r,meas ulated Measured V1 1.12* V2.69* H H H H FT ** ** FT ** ** FT ** ** BB ** ** BB ** ** BB ** ** BB ** ** * Truck-Induced ** Not Available 81

95 Table 4.9 Estimated Yearly Stress Range Histograms Birmingham Structure C ycles per Year S tress R ange (ksi) G age S S V V H H H H H V V FT FT FT FT BT BT BT BT BB BB BB T T

96 Table 4.1 Estimated Yearly Stress Range Historgrams Mobile Structure Cycles per Year Stress Range (ksi) Gage V V H H H H FT FT FT BB BB BB BB

97 Table 4.11 Summary of Measured Effective Stress Range with Comparisons to Measured Maximum and Limit-State Stress Ranges Birmingham Structure Member Type Diagonals & Struts Front Top Chord Back Top Chord Back Bottom Chord WT-Section Gage S re (ksi) S r (ksi) (1 in 1, cycles) S r /S re S r,max (ksi) S r,max /S r S S V V H H H H5.49 * * 1.4 * H V V * * 2.2 * * * 1.97 * * * 2.2 * T T * Not available Table 4.12 Summary of Measured Effective Stress Range with Comparisons to Measured Maximum and Limit-State Stress Ranges Mobile Structure Member Type Gage S re (ksi) S r (ksi) (1 in 1, cycles) S r /S re S r,max (ksi) S r,max /S r V V Diagonals & H Struts H H H FT1.84 * Front Top FT2.88 * Chord FT4.83 * BB1.91 * Back Bottom BB2.85 * Chord BB3.9 * BB4.82 * * Not available 84

98 Table 4.13 Fatigue Life Estimates for all Gage Locations Birmingham Structure Member Type Gage S re,meas (ksi) Cycles/Year (Above 2% of CAFL) Fatigue Life (Years) Diagonals & Struts Front Top Chord Back Top Chord Back Bottom Chord WT-Section S S V V H H H H H V V T T Table 4.14 Fatigue Life Estimates for all Gage Locations Mobile Structure Member Type Diagonals Front Top Chord Back Bottom Chord Gage S re,meas (ksi) Cycles/Year Fatigue Life (Years) (Above 2% of CAFL) V V H H H H FT FT FT BB BB BB BB

99 Back Top Chord (1,2,3,4) Front Top Chord (1,2,3,4) Back Bottom Chord (1,2,3,4) Horiz. Diag. on Top Plane (H3, H4) Horiz. Strut on Bottom Plane (H5, H6) Vert. Diag. on Front Plane (V3, V4) Horiz. Diag. on Bottom Plane (H1, H2) Vert. Diag. at Right Support (V1, V2) WT-Section (T1, T2) Support Diagonal (S1, S2) Figure 4.1 Diagram of Strain Gage Locations for Right Half of Birmingham Structure 86

100 4 in. Strain Gage d chord = in. Fillet Weld Strain Gage Fillet Weld 1 (Top) 2 O.D. Fillet Weld 3 (Front) 4 (Back) 2 O.D. 2 (Bottom) Chord Member Cross Section Strain Gage Typical Strain Gage Location Diagonal Member Figure 4.2. Detail of Strain Gage Placement on Chords and Diagonals 87

101 Strain Gage (T2) 1 in Strain Gage (T1) 1 in.5 in Elevation View Strain Gage (T2) Plan View.25 in Figure 4.3. WT Section Strain Gage Placement on Birmingham Structure 88

102 Horiz. Diag. on Top Plane 1 (H1, H2) Horiz. Diag. on Top Plane 2 (H3, H4) Front Top Chord (1,2,3,4) Back Bottom Chord (1,2,3,4) Vert. Diag. on Front Plane (V1, V2) Figure 4.4. Diagram of Strain Gage Locations on Mobile Structure 89

103 Strain (us) :1. 32:5.3 32:9.6 32: : : : : :35.5 Time (m.s.) Figure 4.5. Truck-Induced Vibration Response of Birmingham Structure from Strain Gage H1 9

104 S tra in (u s ) :1. 32:5.3 32:9.6 32: : : : : :35.5 T im e (m.s. ) Figure 4.6. Truck-Induced Vibration Response of Birmingham Structure fromr Strain Gage FT3 91

105 S tra in (u s ) :1. 32:5.3 32:9.6 32: : : : : :35.5 T im e (m.s. ) Figure 4.7. Truck-Induced Vibration Response of Birmingham Structure from Strain Gage T1 92

106 Strain (us) : : : : :7.2 39: : : : :5.4 Time (m.s.) Figure 4.8. Truck-Induced Vibration Response of Mobile Structure from Strain Gage V1 93

107 Figure 4.9. Pull Down Test Setup 94

108 Figure 4.1. Load Cell and Pulley 95

109 Figure Deflectometer Figure Recording of Data During Pull Down Test 96