FIELD TESTING & LOAD RATING REPORT: MILLARD AVENUE OVER CSX RAILROAD OREGON, OH

Transcription

1 FIELD TESTING & LOAD RATING REPORT: MILLARD AVENUE OVER CSX RAILROAD OREGON, OH SUBMITTED TO: SUBMITTED BY: University of Cincinnati 721 Engineering Research Center, ML30 Cincinnati, OH Phone: (513) BRIDGE DIAGNOSTICS, INC th Court North, Suite 100 Boulder, CO August 2011

2 EXECUTIVE SUMMARY In June of 2011, Bridge Diagnostics, Inc. (BDI) was contracted by the University of Cincinnati to perform diagnostic load testing on two bridges in Toledo and Oregon, Ohio. The purpose of the investigation was to perform a refined analysis and determine if the bridges could carry extremely large permit loads referred to hereafter as superloads, one of which approaches 1,000,000 lbs. The second of the two bridges was Millard Avenue over CSX the railroad (LUC ), which was a 5-span horizontally curved continuous steel beam bridge with a concrete deck. During the field testing phase, the superstructure of the Millard Avenue over CSX was instrumented with a variety of sensors including strain transducers, displacement sensors and inclinometers. Readings were taken from all sensors on a continuous basis while controlled diagnostic load tests were performed. Tests were performed with a single 3-axle dump truck traveling along several lateral paths as well as a single crossing of two dump trucks traveling side by side. Data obtained from the load tests was evaluated for quality and subsequently used to verify and calibrate a finite-element model of the structure. The CSX Bridge was analyzed with a full 3-D finite element model as a means of capturing the complex load response behavior induced by its curved geometry. The goal of the refined analysis was to accurately simulate the structure s complex response to the test loads. This goal was reached by optimizing certain model parameters until an acceptable match between the measured and computed responses was achieved. The calibrated 3-D model was then used to load rate the proposed superloads and determine if these loads could safely cross the bridge according to the AASHTO Standard Specifications for Highway Bridge Design, 17 th Edition and the AASHTO Manual for Condition Evaluation of Bridges, Second Edition. Note that the following assumptions were made during the superload ratings: AASHTO overload considerations were followed in which load factors of 1.3 were applied to both live and dead-load responses while capacities were limited to 0.95 and 0.80 of yield stress to ensure no inelastic deformations in the steel members occurred. The superloads would cross the bridge at crawl speed and the dynamic effects would therefore be negligible. The superloads would travel directly down the center of the bridge while crossing the structure. Pier and foundation components would not be a controlling factor in the load capacity of the bridge. Once the structure was rated for the superloads, the model was also used to load rate all applicable design and legal loads, which included both the Ohio and Michigan legal load configurations. It was found that all components of the superstructure had sufficient capacity to carry all superload configurations with Operating Rating Factors above 1.0. For all load configurations the controlling component was the web and stiffener bearing capacity of the center girder over the piers. The second controlling factor was the girder negative moment capacity at the same locations but these rating factors were significantly larger. For example, the Module-A superload configuration (with no dollies) generated the greatest structural response and in turn II

3 had the lowest ratings of 1.18 for bearing and 1.81 for negative moment. The two variations in the Module-A load configuration (2 and 4 dollies) provided slight improvements of 1.26 and 1.34 for the girder bearing load ratings. The relatively low girder bearing ratings are likely over conservative due to the simplicity of the bearing stiffener design check and may not be applicable considering the limit states applied to the shear ratings. The theoretical failure mode of the girder bearings would only be realistic after the girder webs buckle due to shear. The shear load ratings were based on web buckling (as opposed to post-buckling tension field) with the lowest rating factor of Therefore the basis for the stiffener design will not occur until the Module-A load configuration is essentially doubled. A more realistic capacity calculation would be based on yielding of the web directly over the bearing sole plate; however the design code does not specifically address the situation. Longitudinal stresses in the deck were examined at the negative moment regions over the piers. The live-load stresses due to the Module A loading will approach the tensile strength of the concrete (rupture modulus) indicating that some cracking may occur. It is very likely that the Module A load will induce the largest negative moment ever experienced by this bridge so some level of cracking is to be expected over the piers beyond what is already present. There is no indication that the quantity of additional cracking will be severe. These cracks can be considered superficial from a structural strength or safety view point because at ultimate strength it is always assumed that the concrete cannot resist tension. From a serviceability and maintenance vantage, it would be advisable to apply a sealant to the deck in the negative moment regions after all superload transits. Keeping water and salt from penetrating the deck and corroding the reinforcement would mitigate concerns regarding loss of serviceability life. Inventory Rating Factors generated for all design loads and Ohio and Michigan Legal load configurations were well above 1.0. In general the structure was in very good condition and no load restrictions are required. All pertinent rating results can be found in the following Load Rating Summary Section. This report contains details regarding the instrumentation and load test procedures, a qualitative review of the load test data, a brief explanation of the modeling steps and a summary of the load rating results. Report updates will be made after the first superload crosses the bridge. The first superload transport will generate structural responses much closer to what can be expected by the Module-A superload than did the legal load dump-trucks. If results from the first superload correlate well with the analysis predictions conclusions regarding the potential crossing of Module-A can be obtained relatively quickly. If however, a significant discrepancy between the measurements and predicted responses are obtained, the time required to review the data and modify the analysis could be substantial. Therefore, it is recommended that Module-A be moved later in the transfer schedule. III

4 SUBMITTAL NOTES: This submittal includes the following files on CD: 1. BDI-UT_Superload_Testing_Documents_Mob_1.pdf This file provides pertinent details about the instrumentation plan and testing scenarios/procedures for the first mobilization of the project. 2. BDI-CSX_Curved_Submittal_V1.pdf This is the BDI report in pdf format. It contains details regarding the testing procedures, provides a qualitative data evaluation, displays response histories for each sensor, and discusses any notable observations and/or conclusions arising from the testing process. 3. BDI-CSX_Curved _Input_Files This folder contains all pertinent BDI WinSAC input files that were used in the rating of the structure. See Appendix A for a breakdown of these files. 4. BDI-CSX_Curved _Output_Files This folder contains all pertinent BDI WinSAC output files that were used in the rating of the structure. See Appendix A for a breakdown of these files. 5. BDI-CSX_Curved _Computation_Files This folder contains all pertinent Excel files that were used in the rating of the structure. These files contain the appropriate rating envelopes (non composite dead load, composite dead load, and composite live load) for each critical loading case, and were setup to compute the critical ratings for all the primary rating members. See Appendix A for a breakdown of these files. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH IV

5 TABLE OF CONTENTS EXECUTIVE SUMMARY...II SUBMITTAL NOTES:... IV TABLE OF CONTENTS... V 1. LOAD RATING SUMMARY STRUCTURAL TESTING... 8 GENERAL INFORMATION... 8 TESTING PROCEDURES... 8 PRELIMINARY INVESTIGATION OF TEST RESULTS MODELING, ANALYSIS, AND DATA CORRELATION MODEL CALIBRATION PROCEDURE MODEL CALIBRATION RESULTS LOAD RATING PROCEDURES AND RESULTS RATING PROCEDURES AND ASSUMPTIONS QUALITY CONTROL AND QUALITY ASSURANCE (QC/QA) RATING RESULTS SUMMARY & CONCLUSIONS A. APPENDIX A COMPUTATION INDEX B. APPENDIX B SUPERLOAD CONFIGURATIONS AND WEIGHTS C. APPENDIX C STATE LEGAL LOAD CONFIGURATIONS D. APPENDIX D - REFERENCES FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH V

6 1. LOAD RATING SUMMARY LFR Overload Rating Breakdown TRUCK CONFIGURATION Regenerating Module A No Dollies Regenerating Module A 2 Dollies Regenerating Module A 4 Dollies Regenerating Module B Regenerating Module C Regenerating Module D Lower Stack Section Packinox Exchanger Platforming Reactor Reduction Zone Upper Stack Section Bottom Convection Module Middle Convection Module Top Convection Module POSITIVE MOMENT RF NEGATIVE MOMENT RF SHEAR RF DIAPHRAGM RF BEARING RF CRITICAL RF FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 6

7 LFR Design and Legal Load Critical Rating Summary TRUCK CONFIGURATION HS20 Ohio 2F1 Ohio 3F1 Ohio 4F1 No. 1 No. 2 No. 3 No. 4 No. 5 No. 9 No. 10 No. 11 LIMITING CAPACITY CRITICAL LOCATION Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G Bearing Pier 4, Girder G INVENTORY RATING FACTOR RATING WEIGHT, TONS OPERATING RATING FACTOR RATING WEIGHT, TONS FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 7

8 2. STRUCTURAL TESTING GENERAL INFORMATION The Millard Ave Over CSX Bridge is a five-span, horizontally curved bridge that carries four lanes of Millard Avenue over the CSX railroad interchange in Oregon, OH. The superstructure consists of seven steel plate girders that are composite with a concrete deck. The overall width is approximately 67-0 (48-0 roadway width) and the bridge length is approximately 650 ft. with span lengths varying from 114 ft. to 152 ft. The superstructure s curved geometry was generated in an interesting fashion in that the girders were essentially kinked at the field splices, which were located near the flexural inflection points. The effects of the kinked girder splices were of particular interest in the structural evaluation. Please note that BDI did not perform an in-depth visual inspection of the structure or any non-destructive evaluation (NDE) to verify structural details. TESTING PROCEDURES The structure was instrumented with 114 reusable, surface-mounted strain transducers (Figure Figure 2.5), 3 tiltmeter rotation sensors (Figure 2.6), and 7 Twanger displacement sensors (Figure 2.7) across all five spans. Transducers were installed on both the top and bottom flanges of the girders as well as on selected diaphragms as seen in Figure 2.1 through Figure 2.7. The final instrumentation plans, including sensor locations and IDs, have been provided in the attached drawings labeled UT_Superload_Testing_Documents_Mob1.pdf. Once the instrumentation was installed, a series of diagnostic load tests were completed with the truck traveling across the structure at crawl speed (3 to 5 mph). During testing, data was recorded on all channels at a sample rate of 40 Hz as the Truck 1 test vehicle (70 kip 3-axle dump truck) crossed the structure in the northbound direction along six different lateral positions, referred to as Paths Y1, Y2, Y3, Y4, Y5, and Y6 (further described in the attached testing documents). Additional two-truck tests were conducted with Truck 1 and Truck 2 positioned along paths Y3 and Y2, respectively. These two-truck tests were performed to provide an additional quality check for BDI s finite element model. The truck s longitudinal position was wirelessly tracked so that the response data could later be viewed as a function of vehicle position. During the test procedures, traffic was periodically stopped so that the test truck was the only live load applied to the structure. The entire testing (instrumentation setup, load testing, and tear down) was completed in three days (June th ). Information specific to the load tests can be found in Table 2.1. The test vehicles gross weight, axle weights, and wheel rollout distances (required for tracking its position along the structure) are provided in Table 2.2 and Table 2.3. Test vehicles footprints for Trucks 1 and 2 are also shown in Figure 2.8 and Figure 2.9. The vehicle weights were obtained from certified scales at a local gravel pit, and all vehicle dimensions were measured in the field at the time of testing. BDI would like to thank BP, the University of Toledo, and the University of Cincinnati for their help in scheduling, planning, and organizing the testing project. BDI would also like to thank the Cities of Oregon and Toledo as well as Burkhalter Rigging, Inc. for their excellent field support. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 8

9 Table 2.1 Structure description & testing info. ITEM STRUCTURE NAME BDI Project Number Millard Avenue over CSX OH TESTING DATE June 30, 2011 CLIENT S STRUCTURE ID # LOCATION/ROUTE STRUCTURE TYPE TOTAL NUMBER OF SPANS 5 LUC Millard Avenue Description Steel Superstructure with Composite Deck SPAN LENGTHS Span 1: Span 2: Span 3: Span 4: Span 5: SKEW 0 degrees STRUCTURE/ROADWAY WIDTHS Structure: 67-0 / Roadway: 48-0 WEARING SURFACE OTHER STRUCTURE INFO SPANS TESTED 5 TEST REFERENCE LOCATION (BOW) (X=0,Y=0) TEST VEHICLE DIRECTION TEST BEGINNING POINT LOAD POSITIONS NUMBER/TYPE OF SENSORS SAMPLE RATE NUMBER OF TEST VEHICLES 2 STRUCTURE ACCESS TYPE STRUCTURE ACCESS PROVIDED BY TRAFFIC CONTROL PROVIDED BY Concrete N/A South-west corner of the structure along the inside edge of the curb perpendicular to north curb/expansion joint intersection Westbound Front axle 25.0 ft west of test reference location (BOW) See attached testing documents 114 Strain Transducers, 3 Tiltmeters, 7 Displacement Sensors 40 Hz Snooper, JLG Burkhalter Rigging, Inc. Cities of Oregon and Toledo TOTAL FIELD TESTING TIME 3 days FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 9

10 ITEM Description TEST FILE INFORMATION: FILE NAME LATERAL POSITION FIELD COMMENTS CSX_1 Y1 Good test CSX_2 Y1 Good test. CSX_3 Y2 Good test. CSX_4 Y2 Good test. CSX_5 Y3 Good test. B1308 bad data. 70 kip 3-Axle Dump Truck CSX_6 Y3 Good test. CSX_7 Y4 Good test. CSX_8 Y4 Good test. CSX_9 Y5 Good test. CSX_10 Y5 Good test. CSX_11 Y6 Might have missed first click. CSX_12 Y6 Might have missed first click. Dual Truck Test 70 kip and k 3-Axle Dump Trucks OTHER TEST COMMENTS: CSX_13 Y2 & Y3 Good test. CSX_14 Y2 & Y3 Better test. Weather Warm and Sunny FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 10

11 Figure 2.1 Single surface mounted strain transducer on bottom flange (Typical). Figure 2.2 Transducers on both sides of bottom flange near field splice. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 11

12 Figure 2.3 Surface mounted strain transducer on top flange (Typical). Figure 2.4 Surface mounted strain transducers on top and bottom flange (Typical). FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 12

13 Figure 2.5 Surface mounted strain transducers on diaphragm members (Typical). Figure 2.6 Tiltmeter rotation sensor on bottom flange (Typical). FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 13

14 Figure 2.7 Twanger displacement sensor on bottom flange (Typical). FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 14

15 Table 2.2 Test truck 1 vehicle information. VEHICLE TYPE TRUCK 1 GROSS VEHICLE WEIGHT (GVW) 70,000 lbs WEIGHT/WIDTH - AXLE 1: FRONT 19,500 lbs 7-0 WEIGHT/WIDTH - AXLES 2 & 3: REAR TANDEM PAIR 50,500 lbs (~25,250 lbs each) 7-0 SPACING: AXLE 1 - AXLE SPACING: AXLE 2 AXLE WEIGHTS PROVIDED BY WHEEL ROLLOUT DISTANCE University of Toledo and B.P per wheel revolution Figure 2.8 Test Truck 1 Footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 15

16 Table 2.3 Test truck 2 vehicle information. VEHICLE TYPE TRUCK 1 GROSS VEHICLE WEIGHT (GVW) 66,240 lbs WEIGHT/WIDTH - AXLE 1: FRONT 18,240 lbs 7-2 WEIGHT/WIDTH - AXLES 2 & 3: REAR TANDEM PAIR 48,000 lbs (~24,000 lbs each) 7-2 SPACING: AXLE 1 - AXLE SPACING: AXLE 2 AXLE WEIGHTS PROVIDED BY WHEEL ROLLOUT DISTANCE University of Toledo and B.P. N.A. Figure 2.9 Test Truck 2 Footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 16

17 PRELIMINARY INVESTIGATION OF TEST RESULTS All of the field data was examined graphically to provide a qualitative assessment of the structure's live-load response. Some indicators of data quality include reproducibility between tests along identical truck paths, elastic behavior (strains returning to zero after truck crossing), and any unusual-shaped responses that might indicate nonlinear behavior or possible gage malfunctions. This process can provide a significant amount of insight into how a structure responds to live-load, and is often extremely helpful in performing an efficient and accurate structural analysis. ** - Note that before the test data was reviewed, the strain responses were converted to units of stress assuming an elastic modulus of 29,000 ksi. RESPONSES AS A FUNCTION OF LOAD POSITION: Data recorded from the wireless truck position indicator (BDI AutoClicker) was processed so that the corresponding stress data could be presented as a function of vehicle position. REPRODUCIBILITY AND LINEARITY OF RESPONSES: The structural responses from identical tests were very reproducible as shown in Figure 2.10 through Figure In addition, all responses appeared to be linear with respect to magnitude and truck position, and all stresses returned to nearly zero (within the field accuracy of the sensors), indicating that the structure was acting in a linear-elastic manner. The majority of the response histories had a similar degree of reproducibility and linearity, indicating that the data was of good quality. MAXIMUM STRESSES OBSERVED: The following maximum stresses were observed in each monitored structural component: GIRDERS: A stress of 3.2 ksi was reached in Girder G near midspan of Span 2 under the Truck 1 loading (70 kips) while a stress of 3.8 ksi was recorded in Girder E near midspan of Span 2 under the loading of both Trucks 1 & 2 (136 kips). DIAPHRAGMS: A stress of 2.0 ksi was reached near FS 2 under the Truck 1 loading (70 kips) while a stress of 4.0 ksi was recorded near midspan of Span 2 under the loading of both Trucks 1 & 2 (136 kips). COMPOSITE BEHAVIOR THROUGHOUT ENTIRE STRUCTURE: Measurements taken throughout the structure verified the presence of composite behavior between the deck and the girders. The field-measured neutral axis locations ranged from the upper portion of the girder web to bottom portion of the deck. Figure 2.15 and Figure 2.16 show the top and bottom flange stresses for locations near both midspan and a pier. These responses are consistent with composite behavior and a good indication that the shear studs are performing as intended; therefore, all subsequent load rating calculations were based on composite behavior for positive moment. More importantly, the response data indicated that the structure was also behaving composite near the piers, which is significant since AASHTO does not typically consider composite action from a capacity stand point for this region. OBSERVED LATERAL FLEXURE: An essential parameter to quantify for a horizontally-curved bridge was the lateral flexure induced by its geometry. Figure 2.17 though Figure 2.19 show stress histories from gages on opposite edges of the bottom flanges of an interior girder at Field splice 1 and Field Splice 2 and an exterior girder at Field Splice 1. The difference in magnitude of these gage pairs can be considered twice the magnitude of the lateral flexure in the section. The largest f l / f b ratio recorded under the loading of Truck 1 was approximately 0.31 (i.e., the lateral stress level was about 31% of the average primary bending stress level). FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 17

18 LATERAL LOAD DISTRIBUTION: When evaluating a bridge for the purpose of developing a load rating, the bridge s ability to laterally distribute load is an essential characteristic to quantify. Lateral distribution is most easily observed by plotting the responses from an entire lateral cross-section, as done in Figure 2.20 though Figure The response values shown in these figures correspond to the longitudinal load positions producing the maximum midspan responses for each truck path at Section 2-2 and Section 5-5. From these figures, it can be observed that the structure exhibited a significant level of lateral load distribution across its cross-section. As previously stated, all test data was initially processed and assessed for quality. Then, one set of test data for each truck path was selected for having the best apparent quality. This selected data was then used to calibrate the finite-element (FE) model of the structure, which was in turn used to produce the load ratings. Table 2.4 provides a list of the data files that were used in the FE analysis. Table 2.4 Selected truck path file information. Truck Path Vehicle Type Selected data file Y1 Truck 1 CSX_2.dat Y2 Truck 1 CSX_3.dat Y3 Truck 1 CSX_5.dat Y4 Truck 1 CSX_7.dat Y5 Truck 1 CSX_10.dat Y6 Truck 1 CSX_11.dat Y2 & Y3 Dual Truck CSX_14.dat Path Y3 Midspan Responses Path Y2 Midspan Responses Path Y1 Midspan Responses Figure 2.10 Example of stress response reproducibility for paths Y1-Y3 at Section 2-2. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 18

19 Path Y4 Midspan Responses Path Y5 Midspan Responses Path Y6 Midspan Responses Figure 2.11 Example of stress response reproducibility for paths Y4-Y6 at Section 2-2. Path Y1 Midspan Responses Path Y2 Midspan Responses Path Y3 Midspan Responses Figure 2.12 Example of displacement response reproducibility for paths Y1-Y3 at Section 2-2. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 19

20 Path Y6 Midspan Responses Path Y5 Midspan Responses Path Y4 Midspan Responses Figure 2.13 Example of displacement response reproducibility for paths Y4-Y6 at Section 2-2. Figure 2.14 Example of rotation response reproducibility at Girder E near Pier 1 Path Y3. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 20

21 Gage on Bottom Flange Gage on Top Flange Figure 2.15 Stress response history -Top & bottom flanges - Girder D near midspan of Span 2 (Section 4-4) due to load path Y4. Gage on Top Flange Gage on Bottom Flange Figure 2.16 Stress response history -Top & bottom flanges - Girder E near Pier 1 in Span 1 (Section 3-3) due to load path Y4. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 21

22 Difference between the two stress histories is equal to twice the observed lateral flexure Figure 2.17 Example of measured lateral flexure - Girder F near FS1. Difference between the two stress histories is equal to twice the observed lateral flexure Figure 2.18 Example of measured lateral flexure - Girder G near FS1. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 22

23 Difference between the two stress histories is equal to twice the observed lateral flexure Figure 2.19 Example of measured lateral flexure Girder F near FS2. Figure 2.20 Lateral stress distribution observed in Section 2-2 from Truck 1. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 23

24 Figure 2.21 Lateral displacement distribution observed in Section 2-2 from Truck 1. Figure 2.22 Lateral stress distribution observed in Section 5-5 from Truck 1. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 24

25 3. MODELING, ANALYSIS, AND DATA CORRELATION MODEL CALIBRATION PROCEDURE This section briefly describes the methods and findings of Millard Avenue over CSX Bridge modeling procedures. A list of modeling and analysis parameters specific to this bridge has been provided in Table 3.1 First, geometric data provided in the bridge plans and insight gained from the qualitative data investigation were used to create an initial, three-dimensional finite-element model using BDI s WinGEN modeling software and is illustrated in Figure 3.1. Once the initial model was created, the load test procedures were reproduced using BDI s WinSAC structural analysis and data correlation software. This was done by moving a two-dimensional footprint of the test truck across the model in consecutive load cases that simulated the designated truck path used in the field. The analytical responses of this simulation were then compared to the field responses to validate the model s basic response and to identify any gross modeling deficiencies. Figure 3.1 Finite-element model of superstructure. The goal of the finite element analysis was to obtain a model that could simulate the structural responses obtained during the load test and then provide a basis for load rating the superloads as well as the standard design and legal loads. An iterative process of response comparison and model calibration was performed until an acceptable match was obtained. This process included identification of parameters such as the effective lateral stiffness, continuity of the concrete deck and parapets over the piers, effective end-restraints, and general consistency of behavior at various points. The primary limitation of this analysis and load rating method was that it was based on linear-elastic responses. For the service-level stresses induced by the test loads, this assumption was valid. Additionally, all of the member capacities were based on yield stress or a smaller stress limit controlled by lateral torsional buckling; therefore since the girders capacity were limited to the linear-elastic range the analysis was valid for the rating of the structure. Note that during the rating process, modeling techniques were used to consider the cracking of the concrete over the piers. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 25

26 In the case of this curved structure, the majority of the model calibration effort was spent in simulating the load distribution characteristics of the structure. Based on the review of the test data, the model was made fully composite throughout the structure. The observed endrestraint/pier fixity behavior was modeled using translational axial springs at the bearing locations, while the continuity characteristics were optimized by adjusting the deck stiffness over the piers. Lastly, it was critical to accurately simulate the lateral load distribution provided by the concrete deck and the cross-bracing, along with reproducing the edge stiffness provided by the concrete parapets. Table 3.1 Analysis and model details. - Linear-elastic finite element - stiffness method. ANALYSIS TYPE MODEL GEOMETRY NODAL LOCATIONS MODEL COMPONENTS LIVE-LOAD DEAD-LOAD TOTAL NUMBER OF RESPONSE COMPARISONS MODEL STATISTICS ADJUSTABLE PARAMETERS FOR MODEL CALIBRATION - 3D composed of shell elements, frame elements, and springs (translation and rotation). - Nodes placed at the ends of all frame elements. - Nodes placed at all four corners of each shell element. - Nodes placed at each spring location. - Shell elements for the deck, girder webs, and transitions between deck and girders (i.e., haunch type elements). - Frame elements for the girder flanges, cross-bracing, cross-bracing connections and parapets. - Translational springs representing the vertical support of the superstructure, and the friction-based rotational resistance at the bearing locations. - Boundary conditions representing vertical and transverse restraint at the abutments - 2-D footprints of test trucks consisting of vertical point loads. Truck paths were simulated by series of load cases and the truck footprint moving at 5 ft increments along a curved path. Note that due to the curved geometry of the structure, the design and legal truck footprints were kinked so that the all wheel loads remained along a given lateral path (best approximation of actual load path) - Self-weight of structure. Verification Model (Used to verify model geometry) gage locations x 438 load positions = 45,552 comparisons Calibration Model (Used to calibrate model, without top flange gages) - 68 gage locations x 438 load positions = 29,784 comparisons - 13,560 Nodes - 18,327 Elements Cross-section/Material types Load Cases - 68 Gage locations (Calibration) 1. Slab Stiffness: Primary Slab (E) 2. Slab Stiffness: Intermediate Slab (E) 3. Slab Stiffness: Slab over piers (E) 4. Member Stiffness: Parapet (E) 5. Diaphragm Connection Stiffness: Diaphragm Connection (E) 6. Partial Fixity: At Fixed Pier (Pier 2) (F x & F y ) FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 26

27 MODEL CALIBRATION RESULTS The parameters and model accuracy values resulting from the initial and final bridge models are provided in Table 3.2. Following the optimization procedures, the final model produced correlations with the measured response generated from the Truck 1 loading. This is considered an excellent match for a complex steel structure like the Millard Avenue over CSX Bridge. Table 3.2 Model accuracy & parameter values. MODELING PARAMETER INITIAL MODEL VALUE FINAL MODEL VALUE Slab Stiffness - Primary Slab E [ksi] - Intermediate Slab E [ksi] - Slab over piers E [ksi] Member Stiffness - Parapets E [ksi] Connection Stiffness - Diaphragms E [ksi] Partial Fixity - At Fixed Pier 2 K Fx & K Fy [kip/in] 3,200 3,200 3,200 3,200 29, ,304 3,304 3,304 3, ERROR PARAMETERS INITIAL MODEL VALUE FINAL MODEL VALUE Absolute Error 73, ,922.8 Percent Error 8.6% 4.5% Scale Error 1.6% 0.7% Correlation Coefficient The final model was found to closely match the member stresses as shown in the comparison plots provided in Figure 3.2 though Figure Note that the solid lines in these comparison plots are the recorded response histories while the discrete markers are the computed analytical results. Additionally, the model s midspan lateral distribution of stress closely matched that of the actual structure as shown in Figure 3.20 through Figure The response values shown in these distribution figures correspond to the longitudinal load positions producing the maximum midspan responses for each truck path. Note that the analytical results in the lateral distribution plots are labeled Analysis Path # in the plot legend while the measured responses are labeled by the corresponding data file (e.g., CSX # ). The dual truck load path s analysis results were also compared with the test data results and found to match closely, further validating the model. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 27

28 Deck stiffness: Modeling of R/C elements must take into account the variability of properties such as crack quantity and density, depth of rebar, actual strength of concrete, etc. As a result, the stiffness properties are the effective properties, and may not be the actual material properties. Based on the concrete s design stiffness of approximately 3,800 ksi and the model s calibrated stiffness of 3,304 ksi, as a whole the deck might be considered to be slightly more flexible than excepted. However due to the nature of the optimization process, the effective modulus can also compensate for any dimensional errors or any other factors that might alter the overall stiffness. More important conclusions from this parameter calibration were that the deck stiffness was in a typical range for a concrete deck, the deck was acting composite with the girders over the piers, and the calibrated stiffness value helped produce an accurate longitudinal and lateral load distribution. Parapet stiffness: The resulting parapet modulus was also higher than the original estimated value. This was primarily a result of the approximated modeling of the parapet detail and their effect on the overall stiffness of the bridge. The effective modulus was increased from the original value to in order to help improve the correlation to the structure s lateral load distribution. This was done because the initial model underestimated the additional edge stiffening provided by the parapet. Note that the parapet elements were modeled composite with the structure (i.e., the eccentricity between the parapets and deck was modeled as one-half the slab thickness plus the distance to the neutral axis of the parapet). This parameter calibration was important because it indicated that the parapets actively participated in the resistance of the live-load. Pier Supports: The degree of horizontal stiffness provided by the fixed pier (Pier 2) was determined by comparing the relative magnitudes in negative moment on both sides of the pier as the truck traveled from one span to the other. The resulting pier stiffness that provided the best response comparisons near the pier was a spring value of ~500 kip/in given to each girder. This value was much less than expected for the fixed pier, which indicates that the flexibility of the pier itself allowed the structure to move longitudinally as if the bearing condition was a more of a simple support. This is a very important conclusion because the reduced fixity at Pier 2 significantly influenced the negative moment responses near the pier. Diaphragm Connections: The connection stiffness of the diaphragm elements were optimized in order to simulate the cross-bracings actual contribution to resisting and distributing load. It was found that the diaphragms overall were somewhat rigidly connected to the girder s which was evident by the relatively high magnitude of stress (2.0 ksi under the 70 kip truck) measured in the diaphragms. It was found that the girder stress correlations improved with the partially rigid diaphragm connection as well. Kinked Girder Field Splices: It was observed that the 3-D model did a good job of capturing the lateral flexure that was measured in the bottom flanges near three field splice locations. No extra modeling efforts were required to obtain the desired match but it was found that the lateral flexure component (difference between the pair of bottom flange gages) was highly sensitive to the gage placement along the beam. Further investigation showed that a complete reversal in the lateral flexural was obtained between the staggered diaphragms on each side of the field splices. The lateral flexure was more a function of the diaphragm arrangement than the kinked field splice because the tension in the lower brace members induced an S shape in the girders bottom flanges. Fortunately the field splices are near the girder inflection points so the primary flexural stresses are relatively low. The results do indicate an undesirable consequence to staggered diaphragm placement. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 28

29 Figure 3.2 Final model Stress comparison plot Section 1-1 Paths Y1-Y3. Figure 3.3 Final model - Stress comparison plot Section 1-1 Paths Y4-Y6. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 29

30 Figure 3.4 Final model Stress comparison plot Section 2-2 Paths Y1-Y3. Figure 3.5 Final model - Stress comparison plot Section 2-2 Paths Y4-Y6. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 30

31 Figure 3.6 Final model Displacement comparison plot Section 2-2 Paths Y1-Y3 Figure 3.7 Final model Displacement comparison plot Section 2-2 Paths Y4-Y6 FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 31

32 Figure 3.8 Final model Stress comparison plot Section 3-3 Paths Y1-Y3. Figure 3.9 Final model - Stress comparison plot Section 3-3 Paths Y4-Y6. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 32

33 Figure 3.10 Final model Stress comparison plot Section 4-4 Paths Y1-Y3. Figure 3.11 Final model Stress comparison plot Section 4-4 Paths Y4-Y6. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 33

34 Figure 3.12 Final model Stress comparison plot Section 5-5 Paths Y1-Y3. Figure 3.13 Final model Stress comparison plot Section 5-5 Paths Y4-Y6. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 34

35 Figure 3.14 Final model Stress comparison plot Section 6-6 Paths Y1-Y3. Figure 3.15 Final model Stress comparison plot Section 6-6 Paths Y4-Y6. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 35

36 Figure 3.16 Final model Stress comparison diaphragm gage - B1131. Figure 3.17 Final model Lateral flexure comparison plot for Girder FS1. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 36

37 Figure 3.18 Final model Lateral flexure comparison plot for Girder FS1. Figure 3.19 Final model Lateral flexure comparison plot for Girder FS2. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 37

38 Figure 3.20 Final model Lateral stress 2-2 Truck 1 Edge Paths. Figure 3.21 Final model Lateral stress 2-2 Truck 1 Paths Y2 and Y5. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 38

39 Figure 3.22 Final model Lateral stress 2-2 Truck 1 Center Paths Figure 3.23 Final model Lateral displacement 2-2 Truck 1 Edge Paths. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 39

40 Figure 3.24 Final model Lateral displacement 2-2 Truck 1 Paths Y2 and Y5. Figure 3.25 Final model Lateral displacement 2-2 Truck 1 Center Paths FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 40

41 Figure 3.26 Final model Lateral stress 5-5 Truck 1 Edge Paths. Figure 3.27 Final model Lateral stress 5-5 Truck 1 Paths Y2 and Y5. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 41

42 Figure 3.28 Final model Lateral stress 5-5 Truck 1 Center Paths. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 42

43 4. LOAD RATING PROCEDURES AND RESULTS RATING PROCEDURES AND ASSUMPTIONS Load ratings were performed on all of the appropriate elements in accordance with the AASHTO LFR guidelines. Structural responses were obtained from slightly modified versions of the final calibrated model (hereafter referred to as the rating models), and member capacities were determined using the information found in the provided as-built plans. The rating methods used in BDI s approach closely match typical rating procedures, with the exception that a fieldverified finite-element model analysis was used rather than a typical AASHTO girder-line analysis. This section briefly discusses the methods and findings of the load rating procedures. Four different versions the calibrated model were created in order to simulate the conditions for the following load effect types: construction dead load, superimposed dead load, composite live-load, and non-composite live-load. Construction Dead Load - The model was adjusted so that all the appropriate dead load was induced on the girder sections alone, meaning: The deck stiffnesses were set near zero The end-restraint/ pier-fixity springs were set to zero The parapets weight and stiffness were set to zero The diaphragm stiffness was set to zero (conservative) Composite Live Load This was basically the calibrated model described in the previous section. Reduced parapet stiffness to that of the slab (3200 ksi) The live-load configurations were moved across the structure Non-Composite Live Load The model was adjusted to simulate a cracked deck over the piers. The deck stiffness over the piers was set near zero The parapets stiffness over the piers was reduced to 1000 ksi Reduced the rest of the parapets stiffness to that of the slab (3200 ksi) The live-load configurations were moved across the structure Superimposed Dead Load - The model was adjusted so that the superimposed dead load of the parapets was applied to the structure once the deck was formed. This loading was applied to both the composite and non-composite live-load models, without any stiffness provided to the parapet elements, in order to keep the modeling assumptions consistent. The parapets stiffness was set near zero Parapets self-weight applied to both models (composite/non-composite) Non-composite Model: The deck stiffness over the piers was set near zero FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 43

44 The AASHTO Standard Specifications for Highway Bridge Design, 17 th Edition were used to calculate the following member capacities: Girder shear capacity Girder moment capacity (positive and negative) Girder Bearing capacity Diaphragm axial capacity (tension and compression Transverse slab moment capacity The following assumptions were made during the load rating of the structure: Live-load effects were taken from the both the composite and non-composite FE models and compared with each other. The composite FE model was representative of the structure s current performance as indicated by the load test results and model calibration. The non-composite condition had negligible deck stiffness over the piers which was represented as a cracked deck in the negative moment region. The comparison of these two analysis sets allowed BDI to consider the effects of cracking in the deck on the stresses induced in the girders. This bound approach ensured that the maximum effects for each load configuration were considered. While the Millard Avenue over CSX Bridge has a curved deck, its girders remain straight but follow the curve of the bridge by utilizing kinked field splices. Therefore, the AASHTO Horizontally Curved Steel Girder Highway Bridge Specifications (2003) were not applicable in the rating of this bridge. The lateral flexure effects of the girders at the kinks were localized whereas the code provided stress limitations based on the radius of curvature. Instead the rating procedure considered biaxial flexure from the combined effects of lateral and primary bending. Due to the refined nature of the bridge model, the flexural ratings for the girders were performed in units of flange stress instead of girder cross-section moment. This was done by using the axial and lateral flexure flange stresses from the FE model and a stress limit corresponding to the given section capacity to calculate the superstructure s flexural ratings. It should be noted that due to the structure s composite behavior near midspan and the relatively short un-braced lengths of the compression flanges near the interior supports, the nominal flexural stress limit for most of the girder sections was equal to the yield stress, F y. The diaphragm members were assumed to act as truss members. Therefore, only the pure compression and tension capacity limits and load effects were checked. The shear capacity for the girders was based on a transversely stiffened web based on AASHTO (Eq & Eq ). Even though web stiffener spacing met the requirements specified in , the additional post-buckling shear strength as defined in Eq was not considered in the shear capacity. Due to the exceptionally heavy loads to be applied it was not desirable to consider web buckling as a limit state. Therefore the shear load ratings can be considered a linear-elastic limit state and conservative compared to ultimate strength. The bearing stiffeners were assumed to act as columns in compression along with a portion of the web as specified by AASHTO Additionally, the critical reaction at the interior supports was used as the load effects for ratings of the girder bearing (i.e., it was assumed that the entire support reaction was applied at the top of the section acting FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 44

45 as a column). This design criterion can be considered highly conservative due to the assumption that 100% of the reaction is applied to the top of the bearing column rather than being distributed along the column length (beam depth). The design failure mechanism can be approached only if the girder webs were to buckle due to shear and we were to rely on the webs post-buckling-tension-field action. It is important to note that the shear ratings provided in this report are based on the buckling limit and not post buckling tension field. Therefore ratings based on girder bearing capacity calculations may be overly conservative since they are only applicable after web buckling occurs. As an alternative approach, the compression area of the web and stiffeners directly over the bearing plate was considered. These results in a significantly increased bearing capacity, however the realistic bearing capacity is likely somewhere in between the two values. The field splice capacity was checked and compared to that of the girder capacity in order in ensure that the splices could develop the full strength of the girders. In all cases the field splices were as strong as or stronger than the girder components. None of the girder sections immediately adjacent to the splices were found to be critical because the splices were located near the dead-load inflection points. Since the splices could develop the full strength of the girder sections and the girder sections were not critical the splices were considered adequate and no further ratings were performed. The transverse slab moment capacity was checked for the critical superload, which was Regenerating Module A with no wing dollies. Since the deck rating was more than satisfactory (RF=2.72) for this critical load, deck ratings were not performed for the rest of the load configurations. A summary of the critical girder moment capacity, girder shear capacity, girder bearing axial capacity, diaphragm axial capacity, and transverse slab capacity for the controlling superload case and the design and legal loads have been provided in Table 4.1 through Table 4.4 Table 4.1 Critical girder flange flexural stress limits. ELEMENT Girder D Section 23 Girder D Section 11 Girder G Section 17 Girder G Section 13 FLANGE DIMENSIONS COMPACTNESS LOAD TYPE DIRECTION OF FLEXURE UNBRACED LENGTH Continuously 16 x3/4 Non-compact Positive Super Braced Load 18 x2 1/8 Non-compact Negative 168 STRESS LIMIT 47.5 ksi (0.95F y ) ksi (0.80F y ) Continuously 16 x1 3/4 Non-compact Positive 50 ksi Design Braced Load 18 x1 5/8 Non-compact Negative ksi FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 45

46 Table 4.2 Critical girder shear capacity. LOAD TYPE CRITICAL LOCATION DIMENSIONS SLENDERNESS STIFFENER SPACING, IN. SHEAR CAPACITY, KIPS Super Load Design & Legal Load Girder D - near Pier 2 Girder G - near Pier 3 60 x9/16 Slender k k Table 4.3 Girder Bearing capacity. ELEMENT DIMENSIONS EFFECTIVE WEB WIDTH, IN. WEB HEIGHT, IN. EFFECTIVE BEARING COLUMN AREA, IN 2 STRENGTH LIMIT, KIPS Bearing Stiffener Bearing Base 7 1/4"x7/ in k 7 1/4"x7/8 21 N/A 28.06in k Table 4.4 Diaphragm axial capacities. ELEMENT AISC SHAPE EFFECTIVE TENSILE AREA, IN 2 UNBRACED LENGTH (COMPRESSION), IN. DIRECTION OF LOAD STRENGTH LIMIT, KIPS Diaphragm L3x3x5/ in Tension Compression 89.0 k k Table 4.5 Transverse slab moment capacities. ELEMENT DIRECTION OF MOMENT AREA OF STEEL, IN 2 DEPTH TO STEEL, IN. EQUIVALENT COMPRESSION BLOCK, IN. MOMENT CAPACITY, KIP- IN/IN Deck Positive Negative Load ratings were performed using the rating models according to the AASHTO Manual for Condition Evaluation of Bridges, Second Edition 2003 Revisions (See Table 4.6 for applied rating factors). A single loading condition was considered for each superload where the load was moved directly down the center of the bridge. Load configurations for each superload are provided in Appendix B. Given the 64 wide roadway, one lane loaded through five lanes loaded conditions were considered for the standard design and legal load configurations illustrated in Appendix C. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 46

47 Initially all loading conditions were considered. However, after determining that only the two and three lanes loaded conditions generated the largest responses when considering multilane reduction factors, the other load conditions were neglected in order to reduce the computational time during the analysis. The actual two and three lanes load configurations used have been provided in Figure 4.1 and Figure 4.2, while the critical HS20-44 load configuration is provided in Figure 4.3. The HS-20 lane-load and truck load specified by AASHTO are illustrated in Figure 4.4 Figure 4.5. The centrifugal effects on the live-loads were accounted for using amplification factors that were applied to the inside and outside wheels, which simulated the overturning effect on legal load trucks. Multiple presence factors were applied according to AASHTO All structural dead loads were automatically applied by the modeling program s self-weight function. Table 4.6 Applied LFR Rating factors. FACTOR TYPE DESCRIPTION FACTOR VALUE AASHTO (Overload) AASHTO (Design & Legal Load) Dead Load All 1.3 Live Load All 1.3 Impact Factor All 0 Overload Reduction Non-Composite 0.80 Overload Reduction Composite 0.95 Dead Load - All 1.3 Live Load Inventory 2.17 Live Load - Operating 1.3 Impact Factor All Centrifugal Wheel Amplification Factors (Vertical Loads) Inside Wheel 0.76 Outside Wheel 1.24 FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 47

48 Figure 4.1 Double lane load configurations. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 48

49 Figure 4.2 Three lane load configurations. Figure 4.3 Critical HS20-44 load configuration. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 49

50 Figure 4.4 HS20-44 Lane Load Configuration (AASHTO 2002). Figure 4.5 HS20 Truck Load Configuration (AASHTO 2002) FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 50

51 QUALITY CONTROL AND QUALITY ASSURANCE (QC/QA) In order to explain the different levels of QC/QA undertaken during this refined analysis, the following breakdown has been provided: LOAD TEST DATA REVIEW: All of the data was reviewed by at least two qualified personnel in order to ensure only high quality data was used in the model calibration process, and more importantly to gain an understanding of the required behavior from which an efficient and effective calibration plan could be developed. FINITE-ELEMENT MODEL GENERATION: Based on the provided As-Builts, the full 3D FE model was generated visually, first using CAD software (BrisCAD) which was then imported into WinGEN, BDI s modeling program. When the modeling geometry (nodes), member groups, and elements are imported into WinGEN, the program provides a level of QC/QA by informing the user of any element connectivity errors. Once the model was in WinGEN, member group properties (cross-sectional and material) were defined for each group. These member properties were peer-reviewed along with the overall details of the model at multiple stages of the modeling process. FINITE-ELEMENT MODEL VALIDATION: Once it was determined that the test data was of high quality, the initial FE model was validated by comparing its analytical responses from the simulated truck loadings to those measured in the field. This validation process involved over 45,000 different response comparisons from 104 sensors and over 400 different load cases. FINITE-ELEMENT MODEL CALIBRATION: Once it was determined that the initial model is valid, the model was calibrated based on the reviewed test data. This calibration process was established by the conclusions from the data review and more importantly the rating engineers professional experience. This calibration process included many different iterations and steps, all of which involved over 29,000 different stress comparisons from 68 sensors and over 400 different load cases. Because there are many different solutions to a given problem, this stage of the analysis was heavily dependent on the knowledge and expertise of BDI s rating engineers. CALCULATION OF STRUCTURE S CAPACITY AND LOAD-EFFECTS: Once the model was calibrated, it was used to predict the load effects induced by the rating vehicles. This was done by applying the rating loads to slightly modified versions of the calibrated model. The majority of this process was peer-reviewed including the accuracy of both the loading configurations and the model parameters used in each loading stage. Additionally, the superstructure capacities were calculated using an Excel sheet due to the large number of slight changes in section in built-up girders. These capacity calculations were also peer-reviewed. CALCULATION OF STRUCTURE S CRITICAL RATINGS: Once the structure s capacities and load effects were calculated, Excel files were created that combined the construction deadload, superimposed dead-load, and given live-load effects in order to determine the critically loaded elements under the factored loads. These files were created because different modeling assumptions are necessary at these three loading stages, therefore results from three different models were necessary. These files were by checked for quality multiple times during their creation and utilization by multiple qualified engineers. Additionally, these files were provided to University of Toledo. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 51

52 RATING RESULTS The following is a summary of the load rating factors for the superloads, legal loads, and HS20 design loads developed using the LFR guidelines. The bearing failure condition controlled for all load cases with failure point located along Girder D at Pier 1 for the superloads and along Girder G at Pier 4 for the design and legal loads. Please note that the controlling reactions found at these locations were greater by only 5% or less than those found at similar locations at the other interior supports. The controlling loading condition for the design and legal loads was the three lanes loaded condition along the outside edge of the structure s curve. As discussed in the previous sub-section, the live-load effects generated from both the composite and non-composite models were compared in order to envelope the possible loads induced by the rating loads given that the structure is currently acting composite over the piers. The maximum flexural, shear, and diaphragm axial responses from all load configurations were generated using the non-composite FE live load model, however the controlling bearing response (which was the critical load effect) was generated using the composite FE live load model. These outcomes were expected, since the composite model had much better distribution near the interior supports than non-composite model. For the design ratings, the HS20 truck loading controlled girder positive flexure, girder shear, girder bearing, and diaphragm axial ratings; while the negative moment was controlled by the HS20-44 lane loading. The bridge met rating criteria (RF>1.0) for all rating loads (superloads, design loads, and the legal loads) as shown in Table 4.7 through Table 4.17 with the lowest rating factors resulting from the application of the Regenerating Module A load with no wing dollies assisting in distribution of the load. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 52

53 Table 4.7 LFR superload critical positive moment rating factor and responses. LOAD CONFIGURATION Regenerating Module A No Dollies Regenerating Module A 2 Dollies Regenerating Module A 4 Dollies Regenerating Module B Regenerating Module C Regenerating Module D Lower Stack Section Packinox Exchanger Platforming Reactor Reduction Zone Upper Stack Section Bottom Convection Module Middle Convection Module Top Convection Module CONSTRUCTION DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 Note - Provided responses have not been factored. SUPERIMPOSED DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 LIVE LOAD FLEXURAL STRESS, KIP /INCH 2 CRITICAL RF FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 53

54 Table 4.8 LFR superload critical negative moment rating factor and responses. LOAD CONFIGURATION Regenerating Module A No Dollies Regenerating Module A 2 Dollies Regenerating Module A 4 Dollies Regenerating Module B Regenerating Module C Regenerating Module D Lower Stack Section Packinox Exchanger Platforming Reactor Reduction Zone Upper Stack Section Bottom Convection Module Middle Convection Module Top Convection Module CONSTRUCTION DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 Note - Provided responses have not been factored. SUPERIMPOSED DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 LIVE LOAD FLEXURAL STRESS, KIP /INCH 2 CRITICAL RF FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 54

55 Table 4.9 LFR superload critical shear rating factor and responses. LOAD CONFIGURATION Regenerating Module A No Dollies Regenerating Module A 2 Dollies Regenerating Module A 4 Dollies Regenerating Module B Regenerating Module C Regenerating Module D Lower Stack Section Packinox Exchanger Platforming Reactor Reduction Zone Upper Stack Section Bottom Convection Module Middle Convection Module Top Convection Module CONSTRUCTION DEAD LOAD SHEAR, KIP Note - Provided responses have not been factored. SUPERIMPOSED DEAD LOAD SHEAR, KIP LIVE LOAD SHEAR, KIP CRITICAL RF FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 55

56 Table 4.10 LFR superload critical diaphragm rating factor and responses. LOAD CONFIGURATION Regenerating Module A No Dollies Regenerating Module A 2 Dollies Regenerating Module A 4 Dollies Regenerating Module B Regenerating Module C Regenerating Module D Lower Stack Section Packinox Exchanger Platforming Reactor Reduction Zone Upper Stack Section Bottom Convection Module Middle Convection Module Top Convection Module CONSTRUCTION DEAD LOAD AXIAL FORCE, KIP SUPERIMPOSED DEAD LOAD AXIAL FORCE, KIP LIVE LOAD AXIAL FORCE, KIP CRITICAL RF Note - Provided responses have not been factored. - Due to the unknown staging of construction, it was assumed that the construction dead load effects on the diaphragms was negligible. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 56

57 Table 4.11 LFR superload critical bearing rating factor and responses. LOAD CONFIGURATION Regenerating Module A No Dollies Regenerating Module A 2 Dollies Regenerating Module A 4 Dollies Regenerating Module B Regenerating Module C Regenerating Module D Lower Stack Section Packinox Exchanger Platforming Reactor Reduction Zone Upper Stack Section Bottom Convection Module Middle Convection Module Top Convection Module CONSTRUCTION DEAD LOAD AXIAL FORCE, KIP Note - Provided responses have not been factored. SUPERIMPOSED DEAD LOAD AXIAL FORCE, KIP LIVE LOAD AXIAL FORCE, FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 57 KIP CRITICAL RF Table 4.12 LFR superload critical transverse slab rating factor and responses. LOAD CONFIGURATION Regenerating Module A No Dollies CONSTRUCTION DEAD LOAD MOMENT, KIP-IN/IN Note - Provided responses have not been factored. SUPERIMPOSED DEAD LOAD MOMENT, KIP-IN/IN LIVE LOAD AXIAL FORCE, KIP CRITICAL RF

58 Table 4.13 LFR design/legal load critical positive moment rating factors and responses. LOAD CONFIGURATION CONSTRUCTION DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 COMPOSITE DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 LIVE LOAD FLEXURAL STRESS, KIP /INCH 2 CRITICAL INVENTORY RF CRITICAL OPERATING RF HS Ohio 2F Ohio 3F Ohio 4F No. 1 No. 2 No. 3 No. 4 No. 5 No. 9 No. 10 No. 11 Note: - Provided responses have not been factored. - Live Load responses include impact and centrifugal effects FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 58

59 Table 4.14 LFR design/legal load critical negative moment rating factors and responses. LOAD CONFIGURATION CONSTRUCTION DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 COMPOSITE DEAD LOAD FLEXURAL STRESS, KIP /INCH 2 LIVE LOAD FLEXURAL STRESS, KIP /INCH 2 CRITICAL INVENTORY RF CRITICAL OPERATING RF HS Ohio 2F Ohio 3F Ohio 4F No. 1 No. 2 No. 3 No. 4 No. 5 No. 9 No. 10 No. 11 Note: - Provided responses have not been factored. - Live Load responses include impact and centrifugal effects FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 59

60 Table 4.15 LFR design/legal load critical shear rating factors and responses. LOAD CONFIGURATION CONSTRUCTION DEAD LOAD SHEAR FORCE, KIP COMPOSITE DEAD LOAD SHEAR FORCE, KIP LIVE LOAD SHEAR FORCE, KIP CRITICAL INVENTORY RF CRITICAL OPERATING RF HS Ohio 2F Ohio 3F Ohio 4F No. 1 No. 2 No. 3 No. 4 No. 5 No. 9 No. 10 No. 11 Note: - Provided responses have not been factored. - Live Load responses include impact and centrifugal effects FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 60

61 Table 4.16 LFR design/legal load critical bearing rating factors and responses. LOAD CONFIGURATION CONSTRUCTION DEAD LOAD AXIAL FORCE, KIP COMPOSITE DEAD LOAD AXIAL FORCE, KIP LIVE LOAD AXIAL FORCE, KIP CRITICAL INVENTORY RF CRITICAL OPERATING RF HS Ohio 2F Ohio 3F Ohio 4F No. 1 No. 2 No. 3 No. 4 No. 5 No. 9 No. 10 No. 11 Note: - Provided responses have not been factored. - Live Load responses include impact and centrifugal effects FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 61

62 Table 4.17 LFR design/legal load critical diaphragm rating factors and responses. LOAD CONFIGURATION CONSTRUCTION DEAD LOAD AXIAL FORCE, KIP COMPOSITE DEAD LOAD AXIAL FORCE, KIP LIVE LOAD AXIAL FORCE, KIP CRITICAL INVENTORY RF CRITICAL OPERATING RF HS Ohio 2F Ohio 3F Ohio 4F No. 1 No. 2 No. 3 No. 4 No. 5 No. 9 No. 10 No Note: - Provided responses have not been factored. - Live Load responses include impact and centrifugal effects. - Due to the unknown staging of construction, it was assumed that the construction dead load effects on the diaphragms was negligible. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 62

63 5. SUMMARY & CONCLUSIONS In general, the response data recorded during the load tests was found to be of good quality and indicated no major signs of distress. The following is a summary of the observations made from the preliminary review of the test data: The maximum stress measured under the double truck loading (~136 kips) was ~3.8 to 4.0 ksi, which was measured in both the girders and the diaphragm. Data recorded both near midspan and the piers indicated that the girders were acting composite with the deck throughout the structure. The measured lateral flexure was up to approximately 31% of the average primary bending stress at the kinked field splice locations. The lateral stresses were determined to be localized at the spliced connections and result of the direction change and the staggered diaphragm arrangement near the splices. The Millard Ave over CSX Bridge was analyzed with a full 3-D finite element model as a means of capturing the complex load response behavior induced by its curved geometry. This three-dimensional finite element model was created using the collected structural information, and subsequently calibrated until an acceptable match between the measured and analytical responses was achieved. A very good correlation of between the measured and computed responses was obtained during the modeling process. The resulting calibrated model was then used to analyze the effects of the superloads, and standard AASHTO design and state-based legal loads using the Load Factor Rating (LFR) method. Note that the following assumptions were made during the superload ratings: AASHTO overload considerations were followed in which load factors of 1.3 were applied to both live and dead-load responses while capacities were limited to 0.95 and 0.80 of yield stress to ensure no inelastic deformations in the steel members occurred; The superloads would cross the bridge at crawl speed and the dynamic effects would therefore be negligible. The superloads would travel directly down the center of the bridge while crossing the structure. Pier and foundation components would not be a controlling factor in the load capacity of the bridge. The load rating results indicated that the bridge had satisfactory Operating Load Ratings (RF > 1.0) for all superload configurations. Numerous failure modes were considered including primary flexure of the girders including the effects of biaxial bending at the splice locations, girder shear, girder bearing, diaphragm axial forces and flexural strength of the deck. The controlling failure condition was girder bearing. The critical location for all superloads was located along Girder D at Pier 1. Note that the bearing load ratings at the other piers were very similar. The lowest rating factor was generated under the Regenerating Module A loading (RF=1.18). The corresponding girder moment and shear ratings (1.81 and 2.08) were considerably higher than those obtained for bearing. The bearing ratings can be considered conservative because the theoretical bearing failure mechanism is relevant only after web buckling has occurred due to shear. Overall, the initial results of this project indicate that all of the proposed superloads can safely cross the structure and that there are no significant benefits to the use of dollies to distribute load for the Module A load. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 63

64 Load rating results for all design and legal loads including the Ohio and Michigan truck configuration were satisfactory with Inventory Rating Factors well above 1.0. The girder bearing at Pier 4 along Girder G was the critical location for all legal loads the design loads. The load test, structural investigation, and load rating results presented in this report correspond to the structure at the time of testing. Any structural degradation, damage, and/or retrofits must be taken into account in future ratings. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 64

65 A. APPENDIX A COMPUTATION INDEX The following is a complete list of BDI WinSAC input and output files, along with the example computation Excel sheets that are included on a referenced CD. The rating envelopes (from the output files) for the construction dead load, superimposed dead load, composite live load, and non-composite live load were input into each computation excel file in order compute the critical load factors for each load condition and applicable member capacity. Note that the example computational Excel files can be used to compute a set of ratings for a given load configuration by simply inputting the correct rating envelope data (rto files) into the appropriate spreadsheets. Dead Load Input Files 1. CSX_Rating_ConstDL.inp This BDI WinSAC input file contains all data used in the computing the construction dead load effects on the girders. 2. CSX_Rating_SuperDL_Comp.inp This BDI WinSAC input file contains all data used in the computing the superimposed dead load effects (Parapets) on the girders applied to the composite live-load model. 3. CSX_Rating_SuperDL_NonComp.inp This BDI WinSAC input file contains all data used in the computing the superimposed dead load effects (Parapets) on the girders applied to the non-composite live-load model. Dead Load Rating Output Files 1. CSX_Rating_ConstDL.rto This BDI WinSAC output file contains all data used in the computing the construction dead load effects on the girders. 4. CSX_Rating_SuperDL_Comp.rto This BDI WinSAC output file contains all data used in the computing the superimposed dead load effects (Parapets) on the girders applied to the composite live-load model. 2. CSX_Rating_SuperDL_NonComp.rto This BDI WinSAC output file contains all data used in the computing the superimposed dead load effects (Parapets) on the girders applied to the non-composite live-load model. Superload Live Load Input Files 1. CSX_Rating_LL_modA_ND.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Regenerating Module A with no dollies on the both the composite and non-composite FE models. 2. CSX_Rating_LL_modA_4D.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Regenerating Module A with 4 dollies on the both the composite and non-composite FE models. 3. CSX_Rating_LL_modA_8D.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Regenerating Module A with 8 dollies on the both the composite and non-composite FE models. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 65

66 4. CSX_Rating_LL_modB.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Regenerating Module B on the both the composite and non-composite FE models. 5. CSX_Rating_LL_modC.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Regenerating Module C on the both the composite and non-composite FE models. 6. CSX_Rating_LL_modD.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Regenerating Module D on the both the composite and non-composite FE models. 7. CSX_LL_lowstack.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Lower Stack Section on the both the composite and noncomposite FE models. 8. CSX_Rating_LL_packinox.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Packinox Exchanger on the both the composite and non-composite FE models. 9. CSX_Rating_LL_reactor.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Platforming Reactor on the both the composite and non-composite FE models. 10. CSX_Rating_LL_reduct.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Reduction Zone on the both the composite and noncomposite FE models. 11. CSX_Rating_LL_upstack.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Upper Stack Section on the both the composite and non-composite FE models. 12. CSX_Rating_LL_BOT.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Bottom Convection Module on the both the composite and non-composite FE models. 13. CSX_LL_MID.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Middle Convection Module on the both the composite and noncomposite FE models. 14. CSX_Rating_LL_TOP.inp - This BDI WinSAC input file contains all data used in computing the live load effects of the Top Convection Module on the both the composite and non-composite FE models. Superload Live Load Rating Output Files 1. CSX_Rating_LL_modA_ND.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Regenerating Module A with no dollies on the both the composite and non-composite FE models. 2. CSX_Rating_LL_modA_4D.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Regenerating Module A on the both the composite and non-composite FE models. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 66

67 3. CSX_Rating_LL_modA_8D.rto - This BDI output file contains all envelope data used from the live load effects of the Regenerating Module A with 8 dollies on the both the composite and non-composite FE models. 4. CSX_LL_modB.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Regenerating Module B on the both the composite and noncomposite FE models. 5. CSX_Rating_LL_modC.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Regenerating Module C on the both the composite and non-composite FE models. 6. CSX_Rating_LL_modD.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Regenerating Module D on the both the composite and non-composite FE models. 7. CSX_Rating_LL_lowstack.rto - This BDI output file contains all envelope data used from the live load effects of the Lower Stack Section on the both the composite and noncomposite FE models. 8. CSX_Rating_LL_packinox.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Packinox Exchanger on the both the composite and non-composite FE models. 9. CSX_Rating_LL_reactor.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Platforming Reactor on the both the composite and non-composite FE models. 10. CSX_Rating_LL_reduct.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Reduction Zone on the both the composite and noncomposite FE models. 11. CSX_Rating_LL_upstack.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Upper Stack Section on the both the composite and non-composite FE models. 12. CSX_Rating_LL_BOT.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Bottom Convection Module on the both the composite and non-composite FE models. 13. CSX_Rating_LL_MID.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Middle Convection Module on the both the composite and non-composite FE models. 14. CSX_Rating_LL_TOP.rto - This BDI WinSAC output file contains all envelope data used from the live load effects of the Top Convection Module on the both the composite and non-composite FE models. Design and Legal Live Load Input Files 1. CSX_Rating_HS20_14.inp This BDI WinSAC input file contains all data used in computing the live load effects of the HS20-14 Truck on the both the composite and noncomposite FE models. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 67

68 2. CSX_Rating_Mich01.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 1 on the both the composite and non-composite FE models. 3. CSX_Rating_Mich02.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 2 on the both the composite and non-composite FE models. 4. CSX_Rating_Mich03.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 3 on the both the composite and non-composite FE models. 5. CSX_Rating_Mich04.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 4 on the both the composite and non-composite FE models. 6. CSX_Rating_Mich05.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 5 on the both the composite and non-composite FE models. 7. CSX_Rating_Mich09.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 9 on the both the composite and non-composite FE models. 8. CSX_Rating_Mich10.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 10 on the both the composite and non-composite FE models. 9. CSX_Rating_Mich11.inp This BDI WinSAC input file contains all data used in computing the live load effects of No. 11 on the both the composite and non-composite FE models. 10. CSX_Rating_Ohio2F1.inp This BDI WinSAC input file contains all data used in computing the live load effects of Ohio2F1 Truck on the both the composite and noncomposite FE models. 11. CSX_Rating_Ohio3F1.inp This BDI WinSAC input file contains all data used in computing the live load effects of Ohio3F1 Truck on the both the composite and noncomposite FE models. 12. CSX_Rating_Ohio4F1.inp This BDI WinSAC input file contains all data used in computing the live load effects of Ohio4F1 Truck on the both the composite and noncomposite FE models. Design and Legal Live Load Rating Output Files 1. CSX_Rating_HS20_14.rto This BDI WinSAC output file contains all envelope data used from the live load effects of the HS20-14 Truck on the both the composite and noncomposite FE models. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 68

69 2. CSX_Rating_Mich01.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 1 on the both the composite and noncomposite FE models. 3. CSX_Rating_Mich02.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 2 on the both the composite and noncomposite FE models. 4. CSX_Rating_Mich03.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 3 on the both the composite and noncomposite FE models. 5. CSX_Rating_Mich04.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 4 on the both the composite and noncomposite FE models. 6. CSX_Rating_Mich05.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 5 on the both the composite and noncomposite FE models. 7. CSX_Rating_Mich09.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 9 on the both the composite and noncomposite FE models. 8. CSX_Rating_Mich10.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 10 on the both the composite and noncomposite FE models. 9. CSX_Rating_Mich11.rto This BDI WinSAC output file contains all envelope data used from the live load effects of No. 11 on the both the composite and non-composite FE models. 10. CSX_Rating_Ohio2F1.rto This BDI WinSAC output file contains all envelope data used from the live load effects of Ohio2F1 Truck on the both the composite and noncomposite FE models. 11. CSX_Rating_Ohio3F1.rto This BDI WinSAC output file contains all envelope data used from the live load effects of Ohio3F1 Truck on the both the composite and noncomposite FE models. 12. CSX_RatingOhio4F1.rto This BDI WinSAC output file contains all envelope data used from the live load effects of Ohio4F1 Truck on the both the composite and noncomposite FE models. Please note that all input and output files end in NC for the non-composite model and Comp for the composite model. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 69

70 LFR Critical Computation Excel Files 1. CSX Load Effects-Ratings_modAno_Comp.xlsx This Excel file that was used to compute the critical ratings for the Regenerating Module A with no dollies. The following output file rating envelopes were used in this file: a. CSX_Rating_ConstDL.rto b. CSX_Rating_SuperDL_Comp.rto c. CSX_Rating_LL_C_modA_ND_Comp.rto 2. CSX Load Effects-Ratings_modAno_Comp.xlsx This Excel file that was used to compute the critical ratings for the Regenerating Module A with no dollies. The following output file rating envelopes were used in this file: a. CSX_Rating_ConstDL.rto b. CSX_Rating_SuperDL_NC.rto c. CSX_Rating_LL_C_modA_ND_NC.rto 3. CSX Load Effects-Ratings_HS20_Comp.xlsx This Excel file that was used to compute the critical ratings for the HS20 Loading. The following output file rating envelopes were used in this file: a. CSX_Rating_ConstDL.rto b. CSX_Rating_SuperDL_Comp.rto c. CSX_Rating_HS20_14_Comp.rto 4. CSX Load Effects-Ratings_HS20_NC.xlsx This Excel file that was used to compute the critical ratings for the HS20 Loading. The following output file rating envelopes were used in this file: a. CSX_Rating_ConstDL.rto b. CSX_Rating_SuperDL_NC.rto c. CSX_Rating_HS20_14_NC.rto FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 70

71 B. APPENDIX B SUPERLOAD CONFIGURATIONS AND WEIGHTS Figure B.1 Superload Configurations Regenerating Module A No Dolly and 4 Dolly configuration and weights. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 71

72 Figure B.2 Superload Configurations Regenerating Module A 4 Dolly footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 72

73 Figure B.3 Superload Configurations Regenerating Module A 8 Dolly configuration and weights. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 73

74 Figure B.4 Superload Configurations Regenerating Module A 8 Dolly footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 74

75 Figure B.5 Superload Configurations Regenerating Module B Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 75

76 Figure B.6 Superload Configurations Regenerating Module C Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 76

77 Figure B.7 Superload Configurations Regenerating Module D Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 77

78 Figure B.8 Superload Configurations Lower Stack Section Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 78

79 Figure B.9 Superload Configurations Packinox Exchanger Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 79

80 Figure B.10 Superload Configurations Platforming Reactor Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 80

81 Figure B.11 Superload Configurations Reduction Zone Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 81

82 Figure B.12 Superload Configurations Upper Stack Section Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 82

83 Figure B.13 Superload Configurations Bottom Convection Module Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 83

84 Figure B.14 Superload Configurations Middle Convection Module Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 84

85 Figure B.15 Superload Configurations Top Convection Module Load configuration, weights, and footprint. FIELD TEST AND LOAD RATING REPORT MILLARD AVE OVER CSX - LUC OREGON, OH 85