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Figure 1. Photos. Degradation of conventional piles (Iskander and Hassan 1998). The figure consists of three color photographs representing examples of deteriorated piles. The first photo (A) shows corroded steel piles in a water environment. The second photo (B) shows a degraded concrete pile also in a water environment. The third photograph (C) shows deteriorated timber piles.

Figure 2. Illustration. Common types of composite piles. Shown in this figure are line-drawing cross sections of the three types of composite pile products that are well suited for load-bearing applications. The first type (A) has a steel pipe core surrounded by recycled plastic material. The second type (B) has an inner core of recycled plastic matrix, surrounded by a cage of rebar (FRP or steel). The third type of pile product (C) has an FRP shell filled with nonreinforced concrete.

Figure 3. Graph and Photos. Confinement effect of FRP tube on concrete (Fam and Rizkalla 2001 A, B). This figure presents typical stress-strain curves for (A) the composite stub, (B) the unconfined concrete core, and (C) the FRP shell. An arrow relates each of these curves to a color photograph of the respective materials. The X-axis is strain in units of meter per meter (0.002 to 0.012) and the Y-axis is load in kilonewtons (0 to 700). The chart illustrates the confinement effect of an FRP tube on the concrete core of a composite pile. The graph shows that the capacity of the FRP-concrete composite stub significantly exceeds the summed load-sharing capacity of the FRP and the concrete core. The load strain curve begins to depart from the curve of the unconfined concrete in the vicinity of the unconfined concrete strength. The conversion factors are 1 meter equals 3.28 feet and 1 kilonewton equals 225 poundforce.

Figure 4. Graph. Experimental versus predicted load-strain behavior using Fam and Rizkalla's model. This figure compares the load-strain response predicted by the Fam and Rizkalla variable confinement model with the experimental results obtained for Test Stub Number 1. The X-axis is strain in units of microstrain (positive 15 to negative 20). The Y-axis shows load in kilonewtons (0 to 2500). The chart illustrates that, overall, there is good agreement between the model predictions and experimental data for both lateral or hoop strain and axial or compression strain. At loads below about 1800 kilonewtons, predicted and experimental hoop strain are nearly identical. At loads greater than 1800 kilonewtons, experimentally derived hoop strain is slightly greater than model predictions. Both predicted and experimental axial strain data are nearly identical throughout the entire load range. The conversion factor is 1 kilonewton equals 225 poundforce.

Figure 5. Illustration. Strip elements for sectional analysis (Mirmiran and Shahawy 1996). This figure illustrates the basic concept of the Mirmiran and Shahawy strain compatibility/equilibrium model for predicting short-term flexural capacity. It consists of a simple line drawing representing a cross section of a composite pile. The cross section is divided into a series of strip elements, each containing an FRP element and a concrete element, which are used to integrate the normal stresses over the cross-sectional area. Alongside the cross section is a strain diagram with the following undescribed items listed from top to bottom: epsilon subscript top, epsilon subscript CU, C, Neutral Axis, D subscript core, epsilon subscript bottom, and T subscript FRP.

Figure 6. Graphs. Experimental versus analytical moment-curvature response (adapted from Fam and Rizkalla 2002). This figure consists of two graphs comparing, for Beams 4 and 13, the load-curvature response predicted by the Mirmiran and Shahawy (and Fam) train compatibility/equilibrium model with the experimental load-strain results from Fam and Rizkalla (2002). The X-axis on both charts is curvature in units of the product of 1 over meters times 10 to the minus 3 (0 to 160 and 0 to 30 for Beams 4 and 13, respectively). The Y-axis on both charts is moment in kilonewton-meters (0 to 25 and 0 to 2000 for Beams 4 and 13, respectively). The charts show that there is good agreement between the model predictions (analysis) and the experimental data, particularly with respect to tension stiffening. Specifications for Beams 4 and 13 are included alongside the two charts. The conversion factors are 1 meter equals 3.28 feet and 1 kilonewton-meter equals 737.6 poundforce-feet.

Figure 7. Graph. Interaction diagrams, concrete-filled FRP tubes (Mirmiran 1999). This figure shows the results of a study of combined axial and flexural loads on concrete-filled FRP tubes representing both over reinforcement and under reinforcement. The X-axis is moment measured in kips-inches (0 to 5000) and the Y-axis is axial load measured in kips (0 to 1000), kips being 1000 pound-force (lbf). At similar axial loads of about 100 kips, the moment for the overreinforced specimen (about 4000 kips-inches) is higher than the moment for the underreinforced specimen (about 2000 kips-inches). Conversion factors are 1 kip equals 1,000 poundforce equals 4.45 kilonewtons, and 1 kip-inch equals 0.11303 kilonewton-meter.

Figure 8. Graphs. Moisture absorption-related durability model. This figure presents two graphs that schematically illustrate: (A) the effect of time and temperature on absorption of moisture by FRP composite piles; and (B) the effect of moisture content or time on property values of the FRP composite piles. Increases in both time and temperature contribute to increases in moisture content of the composite pile. Moisture content of the composite pile increases most rapidly during the initial time periods and at the highest temperatures. As time increases, moisture content of the pile levels out. The property value decreases inversely with moisture content of the composite pile or with time.

Figure 9. Images. SEM images showing FRP damage (McBagonluri, et al., 2000). This figure consists of two scanning electron microscope images showing water-related damage to submerged FRP composites. The first image, at times 1020 magnification, and the second image, at times 3000 magnification, each show fiber damage and formation of cracks at the fiber-matrix interface.

Figure 10. Illustration. Influence of soil-pile interface friction on pile capacity. This figure presents a line drawing and two graphs that schematically show the effects of the soil-pile interface on the capacity of the pile. The pile capacity, shaft capacity, and tip capacity are labeled on a simple line drawing of a submerged pile. The ultimate pile capacity is represented as the sum of the shaft capacity and tip capacity. Alongside this drawing are two graphs showing that the soil-pile interface and the friction angle are important factors in determining the ultimate capacity and load transfer characteristics of the pile. The top graph is labeled soil-pile interface shear tests. Its X-axis is labeled capital delta, and its Y-axis is labeled tau. The graph has three lines that rise from left to right to a peak and then level off. The bottom line is labeled sigma subscript N1, the middle line is sigma subscript N2, and the top line is sigma subscript N3. The bottom graph is labeled interface friction angle, small delta. Its X-axis is labeled sigma subscript N and has increments labeled sigma subscript N1, sigma subscript N2, and sigma subscript N3. The Y-axis is labeled tau. Two straight lines begin at the origin and rise from left to right. A horizontal line intersects the uppermost line near the uppermost line's right end; the resulting angle is labeled small delta subscript peak. A second horizontal line intersects the lower line near the lower line's right end; the resulting angle is labeled small delta subscript CV.

Figure 11. Graph. Grain size curves of test sands. This figure is a particle size distribution chart showing grain size properties of the Density and Model test sands. The X-axis is a logarithmic scale and is particle diameter in millimeters (10 to 0.001) (0.39 to 0.000039 inches) and the Y-axis is percentage finer (0 to 100). The grain size of the Density sand ranges from 0.2 to 0.9 millimeters (0.0078 to 0.035 inches) and contains no fines. The Model sand grain size ranges from 0.1 to 2 millimeters (0.0039 to 0.078 inches) and contains about 5 percent fines.

Figure 12. Photos. Microscopic views of the test sands. This figure consists of two photomicrographs of the test sand particles. Each photo includes a 0.5-millimeter (0.02-inch) scale. The first photomicrograph (A) is of the Density sand showing the subrounded to rounded shapes of the grains. The second photomicrograph (B) is of the Model sand showing the subangular to angular shapes of the particles.

Figure 13. Graphs. Direct shear test results for Density sand (average D subscript R equals 70 percent). This figure presents the results of the displacement-controlled direct shear tests. The figure consists of two graphs displaying shear stress-displacement curves (A) and shear stress envelopes (B) for the Density test sand at 70-percent relative density. For the shear stress displacement graph (A), the X-axis is horizontal displacement in millimeters (0 to 14) and the Y-axis is shear stress in kilopascals (0 to 150). At the four normal stresses tested, the sand generally showed peak shear stresses followed by a reduction toward a residual shear stress. The highest peak shear stress, about 140 kilopascals, was observed when the highest normal pressure of 202.3 kilopascals was applied. At the lowest normal pressure of 23.7 kilopascals, the peak shear stress, about 20 kilopascals, was the smallest increase and only slightly above the residual stress. For the shear envelope chart (B), the X-axis is normal stress in kilopascals (0 to 250) and the Y-axis is shear stress in kilopascals (0 to 150). Two solid straight lines and one dashed straight line rise from left to right. The bottom solid line is labeled phi prime subscript CV, the whole term equaling 29.8 degrees. The top solid line is labeled phi prime subscript peak, the whole term equaling 34.7 degrees. The dashed line is labeled nonlinear envelope: phi subscript 0 equals 35.95 degrees, delta phi equals 6.2 degrees. Conversion factors are 1 millimeter equals 0.039 inch and 1 kilopascal equals 0.145 poundforce per square inch.

Figure 14. Graphs. Direct shear test results for Density sand (average D subscript R equals 100 percent). This figure presents the results of the displacement-controlled direct shear tests. The figure consists of two graphs displaying shear stress-displacement curves (A) and shear stress envelopes (B) for the Density test sand at 100-percent relative density. For the shear stress displacement graph (A), the X-axis is horizontal displacement in millimeters (0 to 14) and the Y-axis is shear stress in kilopascals (0 to 175). At the four normal stresses tested, the sand generally showed peak shear stresses followed by a reduction toward a residual shear stress. The highest peak shear stress, about 165 kilopascals, was observed when the highest normal pressure of 202.3 kilopascals was applied. At the lowest normal pressure of 24.8 kilopascals, the peak shear stress, about 22 kilopascals, was the smallest increase and only slightly above the residual stress. For the shear envelope chart (B), the X-axis is normal stress in kilopascals (0 to 250) and the Y-axis is shear stress in kilopascals (0 to 175). Two solid straight lines and one dashed straight line rise from left to right. The bottom solid line is labeled phi prime subscript CV, the whole term equaling 29 degrees. The top solid line is labeled phi prime subscript peak, the whole term equaling 39.3 degrees. The dashed line is labeled nonlinear envelope: phi subscript 0 equals 41.05 degrees, delta phi equals 4.6 degrees. Conversion factors are 1 millimeter equals 0.039 inch and 1 kilopascal equals 0.145 poundforce per square inch.

Figure 15. Graphs. Direct shear test results for Model sand (average D subscript R equals 75 percent). This figure presents the results of the displacement-controlled direct shear tests. The figure consists of two graphs displaying shear stress-displacement curves (A) and shear stress envelopes (B) for the Model test sand at 75-percent relative density. For the shear stress displacement graph (A), the X-axis is horizontal displacement in millimeters (0 to 14) and the Y-axis is shear stress in kilopascals (0 to 150). At the four normal stresses tested, the sand generally showed peak shear stresses followed by a reduction toward a residual shear stress. The highest peak shear stress, about 145 kilopascals, was observed when the highest normal pressure of 156.3 kilopascals was applied. At the lowest normal pressure of 23.7 kilopascals, the peak shear stress, about 30 kilopascals, was the smallest increase and only slightly above the residual stress. For the shear envelope chart (B), the X-axis is normal stress in kilopascals (0 to 200) and the Y-axis is shear stress in kilopascals (0 to 150). Two solid straight lines and one dashed straight line rise from left to right. The bottom solid line is labeled phi prime subscript CV, the whole term equaling 36.2 degrees. The top solid line is labeled phi prime subscript peak, the whole term equaling 43.4 degrees. The dashed line is labeled nonlinear envelope: phi subscript 0 equals 44.6 degrees, delta phi equals 9.8 degrees. Conversion factors are 1 millimeter equals 0.039 inch and 1 kilopascal equals 0.145 poundforce per square inch.

Figure 16. Illustration. Stylus profilometer sketch (Johnson 2000). This figure is a schematic of the profilometer used to measure the surface topography of pile material. From left to right, the figure shows the stylus tip on a sample, the pickup, a drive shaft, a motor drive unit, and a cable to the signal processor.

Figure 17. Chart. Graphical representation of roughness parameters R subscript T, S subscript M, and R subscript A. This figure is a sample tracing of a roughness profile. It consists of peaks and valleys, and illustrates the three roughness parameters typically used to characterize surface topography. The X-axis illustrates the peak widths, identified as capital S subscript small I, and includes four sample lengths (capital S subscript 1 through capital S subscript 4). The Y-axis is capital R subscript small T, defined as the maximum height of the profile or the height of the peak plus the height of the valley. Also shown in red are the reference mean line, small X, and the vertical mean height, small Z. The maximum peak and the maximum valley are each identified on the profile trace. Two formulas are shown below the trace. The mean spacing parameter, capital S subscript small M, is the average width of a peak over the evaluation length. The average roughness, capital R subscript small A, is the quotient of the integral of the absolute value of the roughness profile height divided by the evaluation length, the integral also being the area between the roughness profile and its mean line.

Figure 18. Photo and Graph. Surface characteristics of Lancaster FRP composite pile. This figure consists of a photograph of the Lancaster FRP pile surface and a graph of its corresponding surface roughness profile. The photograph reveals a diagonally linear surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches) and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights varied narrowly between negative 0.05 millimeters (0.002 inches) and positive 0.05 millimeters (0.002 inches).

Figure 19. Photo and Graph. Surface characteristics of Hardcore FRP composite pile. This figure consists of a photograph of the Hardcore FRP pile surface and a graph of its corresponding surface roughness profile. The photograph reveals an irregular and bumpy surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches) and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights varied between negative 0.1 millimeter (0.0039 inches) and about positive 0.07 millimeters (0.0027 inches).

Figure 20. Photo and Graph. Surface characteristics of Hardcore FRP plate. This figure consists of a photograph of the Hardcore FRP plate surface and a graph of its corresponding surface roughness profile. The photograph reveals a faint zigzag surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches) and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights varied narrowly between negative 0.06 millimeters (0.0023 inches) and about positive 0.06 millimeters (0.0023 inches).

Figure 21. Photo and Graph. Surface characteristics of Hardcore surface-treated FRP plate. This figure consists of a photograph of the Hardcore surface-treated FRP plate surface and a graph of its corresponding surface roughness profile. The photograph reveals a regularly bumpy surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches) and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights varied widely between negative 0.17 millimeters (0.0066 inches) and about positive 0.18 millimeters (0.007 inches).

Figure 22. Photo and Graph. Surface characteristics of Plastic Piling plastic composite pile. This figure consists of a photograph of the Plastic composite pile surface and a graph of its corresponding surface roughness profile. The photograph reveals a regularly swirling surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches) and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights varied widely between negative 0.22 millimeters (0.0086 inches) and positive 0.08 millimeters (0.0031 inches).

Figure 23. Photo and Graph. Surface characteristics of prestressed concrete pile. This figure consists of a photograph of the prestressed concrete pile surface and a graph of its corresponding surface roughness profile. The photograph reveals a regularly bumpy surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches) and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights varied widely between negative 0.22 and positive 0.09 millimeters (negative 0.0087 and positive 0.0035 inches).

Figure 24. Photo and Graph. Surface characteristics of steel sheet pile. This figure consists of a photograph of the steel sheet pile surface and a graph of its corresponding surface roughness profile. The photograph reveals a slightly bumpy surface texture, shown with a scale of 5 millimeters (0.2 inches). The X-axis of the surface roughness profile is distance in millimeters (0 to 30) (0 to 1.2 inches), and the Y-axis is height in millimeters (negative 0.25 to positive 0.25) (negative 0.01 inches to positive 0.01 inches). Peaks' heights vary narrowly between negative 0.025 millimeters (0.001 inches) and about positive 0.02 millimeters (0.0008 inches).

Figure 25. Illustration. Sketch of modified interface shear test setup. This figure contains schematics of a modified shear box used to accommodate curved surface materials for interface shear testing. The side (left) and end (right) views both identify the locations of the normal load at the top, the shear load at the sides, the test soil contained within the box, and the FRP shell section. The side view also shows a rigid stratum below the FRP shell. The end view shows the curvature in the top half of the shear box that conforms to the curved pile section.

Figure 26. Graphs. Typical interface shear test results for Density sand (sigma prime subscript N approximately equal to 100 kilopascals). This figure consists of two graphs that present the results of the interface shear tests for the composite (A) and conventional (B) pile test materials using a 60- to 66-percent relative Density sand at a normal stress of about 100 kilopascals. On both graphs, the X-axis is interface displacement in millimeters (0 to 12) and the Y-axis is interface shear stress in kilopascals (0 to 90). Among the composite materials (A), the FRP plate with bonded sand showed the highest peak interface shear stress of about 63 kilopascals. The untreated FRP plate, the Hardcore 24-4 FRP shell, and the PPI plastic showed similar peak interface stress values between 55 and 58 kilopascals. A minimal peak interface stress of about 40 kilopascals was recorded for the Lancaster CP40 FRP shell material. Of the two conventional materials (B), the highest peak interface stress, at about 68 kilopascals, was recorded for the prestressed concrete pile material. Peak interface stress for the steel sheet pile material was about 59 kilopascals. Conversion factors are 1 millimeter equals 0.039 inch and 1 kilopascal equals 0.145 poundforce per square inch.

Figure 27. Graphs. Typical interface shear test results for Model sand (sigma prime subscript N approximately equal to 100 kilopascals). This figure consists of two graphs that present the results of the interface shear tests for the composite (A) and conventional (B) pile test materials using a 60- to 65-percent relative density Model sand at a normal stress of about 100 kilopascals. On both graphs, the X-axis is interface displacement in millimeters (0 to 12) and the Y-axis is interface shear stress in kilopascals (0 to 90). Among the composite materials (A), the FRP plate with bonded sand showed the highest peak interface shear stress of about 80 kilopascals, followed by the PPI plastic at 75 kilopascals, and the untreated FRP plate at about 65 kilopascals. Similar results, 50 to 55 kilopascals, were obtained for the Lancaster CP40 FRP shell material and the Hardcore 24-4 FRP shell material. Of the two conventional materials (B), the highest peak interface stress, at about 72 kilopascals, was recorded for the prestressed concrete pile material. Peak interface stress for the steel sheet pile material was about 63 kilopascals. Conversion factors are 1 millimeter equals 0.039 inch and 1 kilopascal equals 0.145 poundforce per square inch.

Figure 28. Graph. Interface shear strength envelopes for Lancaster Composite FRP shell. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a range between 50 and 110 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript P equals 27.3 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 26 degrees. The lower solid line is for Density sand and is labeled delta subscript P equals 19.7 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 16.6 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 29. Graph. Interface shear strength envelopes for Hardcore Composite FRP shell. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a narrow range between 100 and 110 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript p equals 29.5 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 29.3 degrees. The lower solid line is for Density sand and is labeled delta subscript p equals 29.2 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 27.3 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 30. Graph. Interface shear strength envelopes for untreated Hardcore FRP plate. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a range between 90 and 130 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript P equals 31.7 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 28 degrees. The lower solid line is for Density sand and is labeled delta subscript P equals 28.4 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 25.7 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 31. Graph. Interface shear strength envelopes for treated Hardcore FRP plate. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a range between 110 and 150 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript P equals 37.3 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 32.6 degrees. The lower solid line is for Density sand and is labeled delta subscript P equals 31.9 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 27.8 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 32. Graph. Interface shear strength envelopes for PPI plastic. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a range between 90 and 140 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript P equals 33.4 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 28.8 degrees. The lower solid line is for Density sand and is labeled delta subscript P equals 27.6 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 24.9 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 33. Graph. Interface shear strength envelopes for concrete. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a range between 110 and 150 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript P equals 34.3 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 28.0 degrees. The lower solid line is for Density sand and is labeled delta subscript P equals 33 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 27.7 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 34. Graph. Interface shear strength envelopes for steel. The X-axis of this graph is effective normal stress in kilopascals (0 to 200). The Y-axis is interface shear stress in kilopascals (0 to 150). The graph has four straight lines rising from the origin on the left to a range between 90 and 130 kilopascals on the right. Two of the lines are solid and indicate peak envelope. The remaining two lines are dashed and indicate constant volume (residual) envelope. The upper solid line is for Model sand and is labeled delta subscript p equals 31.2 degrees. The upper dashed line is for Model sand and is labeled delta subscript CV equals 28.6 degrees. The lower solid line is for Density sand and is labeled delta subscript P equals 28.2 degrees. The lower dashed line is for Density sand and is labeled delta subscript CV equals 25.1 degrees. The conversion factor is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 35. Graph. Multiple linear regression on Density sand tan (delta subscript peak) values. The X-axis of this graph is tan (delta subscript peak) from multiple linear regression and ranges from 0.3 to 0.8. The Y-axis is measured tan (delta subscript peak) and ranges from 0.3 to 0.8. The trend line stretches from the origin in the lower left corner to the upper right corner. One of the seven data points is located on the trend line, three are located below it, and three are located above. The coefficient of determination, or R squared, for this fit was 0.662.

Figure 36. Graph. Multiple linear regression on Density sand tan (delta subscript CV) values. The X-axis of this graph is tan (delta subscript CV) from multiple linear regression and ranges from 0.3 to 0.7. The Y-axis is measured tan (delta subscript CV) and ranges from 0.3 to 0.7. The trend line stretches from the origin in the lower left corner to the upper right corner. Three of the seven data points touch the trend line, two are located below it, and two are located above. The coefficient of determination, or R squared, for this fit was 0.581.

Figure 37. Graph. Multiple linear regression on Model sand tan (delta subscript peak) values. The X-axis of this graph is tan (delta subscript peak) from multiple linear regression and ranges from 0.3 to 0.8. The Y-axis is measured tan (delta subscript peak) and ranges from 0.3 to 0.8. The trend line stretches from the origin in the lower left corner to the upper right corner. Four of the seven data points touch the trend line, one is located below it, and two are located above. The coefficient of determination, or R squared, for this fit was 0.892.

Figure 38. Graph. Multiple linear regression on Model sand tan (delta subscript CV) values. The X-axis of this graph is tan (delta subscript CV) from multiple linear regression and ranges from 0.3 to 0.8. The Y-axis is measured tan (delta subscript CV) and ranges from 0.3 to 0.8. The trend line stretches from the origin in the lower left corner to the upper right corner. Two of the seven data points touch the trend line, three are located below it, and two are located above. The coefficient of determination, or R squared, for this fit was 0.474.

Figure 39. Photos. Burnoff testing. This figure consists of two color photographs. The left photograph (A) is the furnace oven used for burnoff testing conducted in this study. The oven is cylindrical in shape and appears to be a laboratory bench-top size. The right photograph (B) is an FRP coupon sample after the resin matrix is burned off in the oven. The photograph shows the diagonal orientation of the fibers and the direction of the longitudinal axis.

Figure 40. Photo. Typical tension test setup. This figure is a color photograph of the setup used to conduct the axial tension tests to evaluate tension properties of the composite materials. The setup consists of a floor-model INSTRON^®test frame, and shows the vertical alignment of the test grips and the load cell.

Figure 41. Photos. Typical split disk test setup. This figure consists of two color photographs of custom-made split disk fixtures used to conduct the hoop tension tests on precision-cut ring specimens of the composite materials. The left photograph (A) shows a closeup of the 61-centimeter (24-inch) split disk fixture containing a sample. The right photograph (B) shows the 30.5-centimeter (12-inch) split disk fixture containing a sample and mounted in the INSTRON^®test frame.

Figure 42. Photo. Freeze-thaw chamber. This figure is a color photograph of the programmable freeze-thaw chamber used to conduct a series of freeze-thaw tests on the composite materials. The figure shows the chamber with the door opened, revealing the stainless steel interior, shelf, and specimen fixture.

Figure 43. Photo and Illustration. Freeze-thaw fixture. This left side (A) of this figure is a color photograph of the freeze-thaw specimen fixture. The fixture is shown with the Lancaster 61-centimeter (24-inch) samples loaded in 4-point bending. On the right side (B) of the figure is a line drawing illustrating, in a cross-section view, 4-point bending of the sample. It shows a slightly bent sample subjected to the fixed 4-point surface strain.

Figure 44. Graph. Average freeze-thaw cycle undergone by FRP samples. This figure illustrates a typical freeze-thaw cycle. The X-axis is time in hours (0 to 3) and the Y-axis is sample temperature in degrees centigrade (negative 20 to positive 10) (negative 4 degrees Fahrenheit to positive 50 degrees Fahrenheit). The figure shows the temperature of the sample, beginning at 0 degrees centigrade (32 degrees Fahrenheit), climbing rapidly to about 6 degrees centigrade (43 degrees Fahrenheit) in 0.3 hours, declining to about negative 17 degrees centigrade (1.4 degrees Fahrenheit) by 2 hours, and rising again to 0 degrees centigrade (32 degrees Fahrenheit) at 2.4 hours, the time for one complete freeze-thaw cycle. Corresponding text on the chart indicates a minimum temperature of negative 17 degrees centigrade (1 degree Fahrenheit) and a maximum temperature of 5.6 degrees centigrade (42 degrees Fahrenheit). One complete cycle consists of a 20-percent thaw and an 80-percent freeze component.

Figure 45. Graph. Representative baseline longitudinal tension stress-strain curves. The figure presents baseline typical axial tension stress-strain curves for the 30.5-centimeter (12-inch) and 61-centimeter (24-inch) Lancaster FRP, and for the 30.5-centimeter (12-inch) and 61-centimeter (24-inch) Hardcore FRP. The X-axis is strain in percent (0 to 2) and the Y-axis is tensile stress in megapascals (0 to 600). The graph indicates that the 30.5-centimeter (12-inch) Hardcore FRP tube exhibited the highest tensile stress, rising to approximately 475 megapascals at a strain of 1.8 percent. The 61-centimeter (24-inch) Lancaster FRP showed the lowest tensile stress, rising to approximately 100 megapascals at a strain of 1 percent. The 61-centimeter (24-inch) Hardcore FRP rose to approximately 425 megapascals at a strain of 1.8 percent, and the 30.5-centimeter (12-inch) Lancaster FRP rose to a peak of approximately 250 megapascals at a strain of 1.5 percent. One megapascal equals 145 pounds-force per square inch.

Figure 46. Graph. Representative baseline hoop tension stress-strain curves. The figure presents baseline hoop tension stress-strain curves for the composite materials determined by the split disk tests. Measurements were made on circular samples of the 30.5-centimeter (12-inch) and 61-centimeter (24-inch) Lancaster FRP and on the 30.5-centimeter (12-inch) Hardcore FRP materials. The X-axis is strain in percent (0 to 1.6) and the Y-axis is tensile stress in megapascals (0 to 250). The chart indicates that the 30.5-centimeter (12-inch) Lancaster FRP material exhibited the highest hoop tensile stress, rising to approximately 190 megapascals at a strain of just under 1.4 percent. The 30.5-centimeter (12-inch) Hardcore FRP material had the lowest tensile stress, rising to approximately 90 megapascals at a strain of just over 1.2 percent. The tensile stress of the 61-centimeter (24-inch) Lancaster FRP materials rose to approximately 120 megapascals at a strain of just under 1.4 percent.

Figure 47. Graph. Absorption curves for Lancaster 12-inch FRP tube. This figure presents the changes over time in the moisture absorption by samples of the 30.5-centimeter (12-inch)Lancaster FRP material at six different water temperatures. The X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 140). The Y-axis is moisture content in percent (0 to 1.2). The water temperatures of the tests are 22, 35, 45, 55, 65, and 80 degrees centigrade (72, 95, 113, 131, 149, and 176 degrees Fahrenheit). Six curves, one for each water temperature, are shown on the graph. Except for the sample immersed in 80-degree-centigrade (176-degree-Fahrenheit) water, the moisture content of the samples increases rapidly near the beginning of the tests and generally stabilizes. At all five test temperatures, except the highest, the maximum moisture content of the samples is near or below 0.4 percent after 577 days. For the test conducted at 80 degrees centigrade (176 degrees Fahrenheit), maximum moisture content of the sample increases dramatically to about 1 percent at 100 square root hours. In general, the results indicate that moisture absorption by the 30.5-centimeter (12-inch)Lancaster samples increases as the water temperature increases.

Figure 48. Absorption curves for Lancaster 24-inch FRP tube. Chart. This figure presents the changes over time in the moisture absorption by samples of the 61-centimeter (24-inch)Lancaster FRP material at six different water temperatures. The X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 160). The Y-axis is moisture content in percent (0 to 1.6). The water temperatures of the tests are 22, 35, 45, 55, 65, and 80 degrees centigrade (72, 95, 113, 131, 149, and 176 degrees Fahrenheit). Six curves, one for each water temperature, are shown on the graph. The moisture content of the samples at the four lowest temperatures increases rapidly near the beginning of the tests and generally stabilizes near the maximum. After 819 days, the moisture content of the samples at these four lowest temperatures is below about 0.6 percent. For the test conducted at 65 degrees centigrade (149 degrees Fahrenheit), maximum moisture content of the sample increases dramatically to a maximum of almost 1.2 percent and does not appear to level off. Moisture content of the sample tested at 80 degrees centigrade (176 degrees Fahrenheit) increases rapidly, peaks at about 1.5 percent at approximately 60 square root hours, and then sharply decreases to nearly zero.

Figure 49. Graph. Absorption curves for Hardcore 12-inch FRP tube. This figure presents the changes over time in the moisture absorption by samples of the 30.5-centimeter (12-inch) Hardcore FRP material at six different water temperatures. The X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 140). The Y-axis is moisture content in percent (0 to 1.8). The water temperatures of the tests are 22, 35, 45, 55, 65, and 80 degrees centigrade (72, 95, 113, 131, 149, and 176 degrees Fahrenheit). Six curves, one for each water temperature, are shown on the graph. For the five lowest temperature, the moisture content of the samples increases rapidly near the beginning of the tests and generally stabilizes near the maximum. At these five test temperatures, the maximum moisture content of the samples is near or below about 0.25 percent after 578 days. For the test conducted at 80 degrees centigrade (176 degrees Fahrenheit), moisture content of the Hardcore sample increases dramatically to a maximum of about 1.7 percent and shows no sign of leveling off..

Figure 50. Graph. Absorption curves for Hardcore 24-inch FRP tube. This figure presents the changes over time in the moisture absorption by samples of the 61-centimeter (24-inch) Hardcore FRP material at six different water temperatures. The X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 160). The Y-axis is moisture content in percent (0 to 0.5). The water temperatures of the tests are 22, 35, 45, 55, 65, and 80 degrees centigrade (72, 95, 113, 131, 149, and 176 degrees Fahrenheit). Six curves, one for each water temperature, are shown on the graph. The absorption curves are similar for the samples tested at the five lowest water temperatures. The moisture content for these five groups of samples increases rapidly near the beginning of the tests and generally stabilizes. The moisture content of the samples immersed at these temperatures is between 0.2 and 0.3 percent after 681 days. The moisture content of the sample immersed in 80-degree-centigrade (176-degree-Fahrenheit) water reaches a maximum of about 0.45 percent and shows no sign of leveling off.

Figure 51. Graphs. Selected diffusion analyses for Lancaster 12-inch FRP samples. This figure consists of three graphs, each representing a unique test temperature and two curve fits of the experimentally obtained absorption data for the Lancaster 30.5-centimeter (12-inch) FRP samples using both the Fickian diffusion model and the Langmuirian diffusion model. The top graph (A) presents data for the 22-degree-centigrade (72-degree-Fahrenheit) water; the middle graph (B) presents the 45-degree-centigrade (113-degree-Fahrenheit) water test results; and the bottom graph (C) shows the 65-degree-centigrade (149-degree-Fahrenheit) data. On all graphs, the X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 140). The Y-axis is moisture content in percent (0 to 0.4 or 0.5). The experimental data fit both models well in the initial absorption period and near the maximum absorption at the end of the test. For moisture absorption during the middle of the test period, all three graphs show that the Langmuirian model tracks the experimental absorption data more closely than the Fickian model.

Figure 52. Graphs. Selected diffusion analyses for Lancaster 24-inch FRP samples. This figure consists of three graphs, each representing a unique test temperature and two curve fits of the experimentally obtained absorption data for the Lancaster 61-centimeter (24-inch) FRP samples using both the Fickian diffusion model and the Langmuirian diffusion model. The top graph (A) presents data for the 22-degree-centigrade (72-degree-Fahrenheit) water; the middle graph (B) presents the 45-degree-centigrade (113-degree-Fahrenheit) water test results; and the bottom graph (C) shows the 65-degree-centigrade (149-degree-Fahrenheit) data. On all graphs, the X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 160; 0 to 160; and 0 to 180). The Y-axis is moisture content in percent (0 to 0.6; 0 to 0.5; and 0 to 1.4). For the 22-degree-centigrade (72-degree-Fahrenheit) test (graph A) and the 45-degree-centigrade (113-degree-Fahrenheit) test (graph B), the experimental data fit both models well in the initial absorption period and near the maximum absorption at the end of the test. For moisture absorption during the middle of the test period, graph A and graph B show that the Langmuirian model tracks the experimental absorption data more closely than the Fickian model. The 65-degree-centigrade (149-degree-Fahrenheit) test data (graph C) appears to be tracked more closely by the Langmuirian model during the duration of the test.

Figure 53. Graphs. Selected diffusion analyses for Hardcore 12-inch FRP samples. This figure consists of three graphs, each representing a unique test temperature and two curve fits of the experimentally obtained absorption data for the Hardcore 30.5-centimeter (12-inch) FRP samples using both the Fickian diffusion model and the Langmuirian diffusion model. The top graph (A) presents data for the 22-degree-centigrade (72-degree-Fahrenheit) water; the middle graph (B) presents the 45-degree-centigrade (113-degree-Fahrenheit) water test results; and the bottom graph (C) shows the 65-degree-centigrade (149-degree-Fahrenheit) data. On all graphs, the X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 140). The Y-axis is moisture content in percent (0 to 0.25 or 0.30). The curves from the Langmuirian and Fickian models are very similar in the three charts, and both models track the experimental data fairly well and fairly similarly, although the 45-degree-centigrade (113-degree-Fahrenheit) data (graph B) are slightly more aligned with the Langmuirian model curve.

Figure 54. Graphs. Selected diffusion analyses for Hardcore 24-inch FRP samples. This figure consists of three graphs, each representing a unique test temperature and two curve fits of the experimentally obtained absorption data for the Hardcore 61-centimeter (24-inch) FRP samples using both the Fickian diffusion model and the Langmuirian diffusion model. The top graph (A) presents data for the 22-degree-centigrade (72-degrees Fahrenheit) water; the middle graph (B) presents the 45-degree-centigrade (113-degree-Fahrenheit) water test results; and the bottom graph (C) shows the 65-degree-centigrade (149-degree-Fahrenheit) data. On all graphs, the X-axis is the square root of time, time being measured in hours; thus the X-axis is hours to the one-half power (0 to 140). The Y-axis is moisture content in percent (0 to 0.3 or 0.25). The experimental data fit both models well in the initial absorption period and near the maximum absorption at the end of the test. For moisture absorption during the middle of the test period, all three graphs show that the Langmuirian model tracks the experimental absorption data more closely than the Fickian model.

Figure 55. Graph. Longitudinal tensile properties versus submergence time for Lancaster 24-inch FRP tube. This figure is a bar graph presenting the variations in three longitudinal tensile properties of the 61-centimeter (24-inch) Lancaster FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 4, 10, 26, 66, 185, 444, and 787) and the Y-axis is the normalized value of the tensile property (0 to 1.4). For each of the eight submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is considerable scatter in the data sets. Based on the longitudinal tensile property data obtained over the period of submergence, no significant degradation is indicated.

Figure 56. Graph. Hoop tensile properties versus submergence time for Lancaster 24-inch FRP tube. This figure is a bar graph presenting the variations in three hoop tensile properties of the 61-centimeter (24-inch) Lancaster FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 55, 127, 252, 391, and 525) and the Y-axis is the normalized value of the tensile property (0 to 1.4). For each of the six submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is considerable scatter in the data sets. Over the 525-day period of submergence, the hoop tensile strength showed the least change, decreasing by only 8 percent. Initial stiffness and strain to failure revealed greater degradation, but not to a significant extent.

Figure 57. Graph. Longitudinal tensile properties versus submergence time for Lancaster 12-inch FRP tube. This figure is a bar graph presenting the variations in three longitudinal tensile properties of the 30.5-centimeter (12-inch) Lancaster FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 56, 116, 253, and 399) and the Y-axis is the normalized value of the tensile property (0 to 1.4). For each of the five submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is some scatter in the data sets, but less than that observed for the 61-centimeter (24-inch)Lancaster samples. After nearly 400 days of submergence, the longitudinal tensile property showing the greatest degradation is the strain to failure, which decreased nearly 20 percent in the 30.5-centimeter (12-inch)Lancaster sample. The data indicate a smaller decline in tensile strength over the same time period. Little variation in the initial stiffness parameter is shown.

Figure 58. Graph. Hoop tensile properties versus submergence time for Lancaster 12-inch FRP tube. This figure is a bar graph presenting the variations in three hoop tensile properties of the 30.5-centimeter (12-inch) Lancaster FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 102, 254, 384, and 453) and the Y-axis is the normalized value of the tensile property (0 to 1.4). For each of the five submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is some scatter in the data sets. Over the 453-day period of submergence, the hoop tensile strength and initial stiffness both decreased by about 18 percent, significantly more than observed for the 61-centimeter (24-inch)Lancaster sample. The strain-to-failure parameter revealed the least amount of variation, remaining relatively constant over time.

Figure 59. Charts. Longitudinal tensile properties versus moisture content for Lancaster 12-inch FRP tube. This figure consists of two scatter charts illustrating the effect of moisture content on longitudinal tensile strength and stiffness of the 30.5-centimeter (12-inch) Lancaster FRP sample. The top chart (A) presents the variation in tensile strength with increases in moisture content and the bottom chart (B) presents the variation in stiffness with increases in moisture content. The X-axis on both charts is moisture content in percent (0 to 0.6), and the Y-axis is the normalized value of each parameter (0 to 1.2 and 0 to 1.4). Individual data points, representing normalized values for each parameter, are shown primarily for moisture contents between 0.2 and 0.5 percent, corresponding to about 26 and 420 days submergence, respectively. Chart A indicates that longitudinal tensile strength of the 30.5-centimeter (12-inch) Lancaster FRP sample declines gradually as the moisture content increases, but at the higher end of the moisture scale, the tensile strength stabilizes. Chart B indicates that this trend is not evident for the longitudinal stiffness parameter. Between 0.2 and 0.5 percent moisture, stiffness data points are scattered irregularly around the baseline value of 1.

Figure 60. Charts. Hoop tensile properties versus moisture content for Lancaster 12-inch FRP tube. This figure consists of two scatter charts illustrating the effect of moisture content on hoop tensile strength and stiffness of the 30.5-centimeter (12-inch) Lancaster FRP sample. The top chart (A) presents the variation in tensile strength with increases in moisture content and the bottom chart (B) presents the variation in stiffness with increases in moisture content. The X-axis on both charts is moisture content in percent (0 to 0.8), and the Y-axis is the normalized value of each parameter (0 to 1.2 and 0 to 1.4). Individual data points, representing normalized values for each parameter, are shown across the full moisture content scale. Chart A indicates that hoop tensile strength of the 30.5-centimeter (12-inch) Lancaster FRP sample declines gradually as the moisture content increases, but at the higher end of the moisture scale (0.6 to 0.8 percent), this decline appears to stabilize. The data points on chart B indicate a similar behavior, showing a gradual decline in the stiffness with increased moisture content. At moisture levels greater than about 0.4 percent, stiffness values for the 30.5-centimeter (12-inch)Lancaster sample appear to stabilize.

Figure 61. Graph. Longitudinal tensile properties versus submergence time for Hardcore 24-inch FRP tube. This figure is a bar graph presenting the variations in three longitudinal tensile properties of the 61-centimeter (24-inch) Hardcore FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 71, 111, 161, and 636), and the Y-axis is the normalized value of the tensile property (0 to 1.4). For each of the five submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is considerable scatter in the initial stiffness and the strain to failure data sets. The bar graph indicates that, after 636 days of submergence, the longitudinal tensile strength declined by about 13 percent and the strain-to-failure parameter decreased by about 20 percent. No trend for initial stiffness is apparent.

Figure 62. Charts. Longitudinal tensile properties versus moisture content for Hardcore 24-inch FRP tube. This figure consists of two scatter charts illustrating the effect of moisture content on longitudinal tensile strength and initial stiffness of the 61-centimeter (24-inch) Hardcore FRP sample. The top chart (A) presents the variation in tensile strength with increases in moisture content and the bottom chart (B) presents the variation in stiffness with increases in moisture content. The X-axis on both charts is moisture content in percent (0 to 0.3), and the Y-axis is the normalized value of each parameter (0 to 1.4 and 0 to 1.6). Individual data points, representing normalized values for each parameter, are shown primarily for moisture contents between 0.1 and 0.3 percent. Chart A indicates that longitudinal tensile strength of the 61-centimeter (24-inch) Hardcore FRP sample declines gradually as the moisture content increases. The data show a 16-percent decline in longitudinal tensile strength at the highest moisture content of 0.33 percent. The stiffness data points on chart B are widely scattered around the baseline and do not suggest any obvious trends.

Figure 63. Graph. Longitudinal tensile properties versus submergence time for Hardcore 12-inch FRP tube. This figure is a bar graph presenting the variations in three longitudinal tensile properties of the 30.5-centimeter (12-inch) Hardcore FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 42, 115, 252, and 398) and the Y-axis is the normalized value of the tensile property (0 to 1.6). For each of the five submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is considerable scatter in the data sets. After nearly 400 days of submergence, the longitudinal tensile strength and the strain to failure parameters for the 30.5-centimeter (12-inch) Hardcore sample each declined by approximately 25 percent. Over the same time period, the initial stiffness increased by about 10 percent.

Figure 64. Graph. Hoop tensile properties versus submergence time for Hardcore 12-inch FRP tube. This figure is a bar graph presenting the variations in three hoop tensile properties of the 30.5-centimeter (12-inch) Hardcore FRP sample with time submerged in the water. The X-axis is submergence time in days (0, 118, 265, 401, and 464) and the Y-axis is the normalized value of the tensile property (0 to 2). For each of the five submergence times, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. Over the 464-day period of submergence, the data indicate that the hoop tensile strength decreased by about 20 percent. The error bars reveal relatively high scatter among the data sets for initial stiffness and strain to failure, hindering the identification of any trends over time.

Figure 65. Charts. Longitudinal tensile properties versus moisture content for Hardcore 12-inch FRP tube. This figure consists of two scatter charts illustrating the effect of moisture content on longitudinal tensile strength and stiffness of the 30.5-centimeter (12-inch) Hardcore FRP sample. The top chart (A) presents the variation in tensile strength with increases in moisture content and the bottom chart (B) presents the variation in stiffness with increases in moisture content. The X-axis on both charts is moisture content in percent (0 to 0.35) and the Y-axis is the normalized value of each parameter (0 to 1.2 and 0 to 1.8). Individual data points, representing normalized values for each parameter, are shown primarily for moisture contents between 0.1 and 0.3 percent. Chart A indicates that longitudinal tensile strength of the 30.5-centimeter (12-inch) Hardcore FRP sample declines as the moisture content increases. A 25-percent maximum degradation in tensile strength occurs at a moisture content of about 0.25 percent. The stiffness data points on chart B are widely scattered primarily above the baseline and do not suggest any obvious trends.

Figure 66. Charts. Hoop tensile properties versus moisture content for Hardcore 12-inch FRP tube. This figure consists of two scatter charts illustrating the effect of moisture content on hoop tensile strength and stiffness of the 30.5-centimeter (12-inch) Hardcore FRP sample. The top chart (A) presents the variation in tensile strength with increases in moisture content and the bottom chart (B) presents the variation in stiffness with increases in moisture content. The X-axis on both charts is moisture content in percent (0 to 0.8) and the Y-axis is the normalized value of each parameter (0 to 1.4). Individual data points, representing normalized values for each parameter, are shown primarily for moisture contents between 0.2 and 0.7 percent. The widely scattered data sets for hoop tensile strength and stiffness shown in charts A and B, respectively, preclude identification of any trends.

Figure 67. Graph. Influence of freeze-thaw cycling on the longitudinal tensile properties for the Lancaster 24-inch FRP tube. This figure is a bar graph presenting variations in three longitudinal tensile properties of the 61-centimeter (24-inch) Lancaster FRP sample with freeze-thaw cycles. The X-axis is the number of freeze-thaw cycles (0, 100, 300, and 500) and the Y-axis is the normalized value of the tensile property (0 to 1.5). For each of the four freeze-thaw cycles, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is some scatter in the data sets. The most notable change after 500 freeze-thaw cycles is the longitudinal tensile strength, which declined by about 25 percent. Taking the standard deviations into account, the variation in the initial stiffness and the strain to failure parameters is minor.

Figure 68. Graph. Influence of freeze-thaw cycling on the longitudinal tensile properties for the Lancaster 12-inch FRP tube. This figure is a bar graph presenting variations in three longitudinal tensile properties of the 30.5-centimeter (12-inch) Lancaster FRP sample with freeze-thaw cycles. The X-axis is the number of freeze-thaw cycles (0, 100, 300, and 500) and the Y-axis is the normalized value of the tensile property (0 to 1.75). For each of the four freeze-thaw cycles, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is significant scatter in the data sets. Taking the data scatter into account, the variations in the three tensile properties over 500 freeze-thaw cycles appear to be minor.

Figure 69. Graph. Influence of freeze-thaw cycling on the longitudinal tensile properties for the Hardcore 24-inch FRP tube. This figure is a bar graph presenting variations in three longitudinal tensile properties of the 61-centimeter (24-inch) Hardcore FRP sample with freeze-thaw cycles. The X-axis is the number of freeze-thaw cycles (0, 100, 300, and 500) and the Y-axis is the normalized value of the tensile property (0 to 1.5). For each of the four freeze-thaw cycles, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is some scatter in the data sets. Taking the data scatter into account, the variations in the three tensile properties over 500 freeze-thaw cycles appear to be minor.

Figure 70. Graph. Influence of freeze-thaw cycling on the longitudinal tensile properties for the Hardcore 12-inch FRP tube. This figure is a bar graph presenting variations in three longitudinal tensile properties of the 30.5-centimeter (12-inch) Hardcore FRP sample with freeze-thaw cycles. The X-axis is the number of freeze-thaw cycles (0 and 500) and the Y-axis is the normalized value of the tensile property (0 to 1.5). For each of the two freeze-thaw cycles, three normalized vertical property bars are given: one for strength, one for initial stiffness, and one for strain to failure. Error bars for plus and minus one standard deviation are shown on each of the property bars. The error bar brackets reveal that there is some scatter, primarily in the strain-to-failure data set. The bar chart does not show degradation in any of the three longitudinal tensile properties of the 30.5-centimeter (12-inch) Hardcore sample after 500 freeze-thaw cycles.

Figure 71. Photo. SEM images of Lancaster 24-inch FRP tube. This figure consists of three black-and-white scanning electron micrographs of the 61-centimeter (24-inch) Lancaster FRP material. The first micrograph (A) on the left is the material in the original condition it was received. The center micrograph (B) is a sample at moisture saturation, and the sample shown on the right (C) is after freeze-thaw cycling. All micrographs are magnified times 1000. In the condition received, as shown in micrograph A, the micrograph shows densely packed circular components with entire margins. At moisture saturation (B), the micrograph shows circular components with some fraying at the margins. After freeze-thaw cycles (C), the micrograph appears similar to that of the original condition and consists of circular components showing no sign of fraying at the margins.

Figure 72. Photo. SEM images of Lancaster 12-inch FRP tube. This figure consists of three black-and-white scanning electron micrographs of the 30.5-centimeter (12-inch) Lancaster FRP material. The first micrograph (A) on the left is the material in the original condition it was received. The center micrograph (B) is a sample at moisture saturation, and the sample shown on the right (C) is after freeze-thaw cycling. All micrographs are magnified times 1000. In the condition received (A), the micrograph shows densely packed circular components with entire margins and some variation in size. At moisture saturation (B), the micrograph is similar to that of the original condition with no evidence of damage. After the freeze-thaw cycles (C), the micrograph appears similar to the original and consists of densely packed circular components of varying size. No evidence of damage is visible.

Figure 73. Photo. SEM images of Hardcore 24-inch FRP tube. This figure consists of three black-and-white scanning electron micrographs of the 61-centimeter (24-inch) Hardcore FRP material. The first micrograph (A) on the left is the material in the original condition it was received. The center micrograph (B) is a sample at moisture saturation, and the sample shown on the right (C) is after freeze-thaw cycling. All micrographs are magnified times 1000. In the condition received (A), the micrograph shows circular components with entire margins and some variation in size. At moisture saturation (B), the micrograph consists of circular components with some evidence of fraying at the margins. After the freeze-thaw cycles (C), the micrograph appears similar to that of the original material and consists of densely packed circular components of varying size. No evidence of damage is visible.

Figure 74. Photo. SEM images of Hardcore 12-inch FRP tube. This figure consists of three black-and-white scanning electron micrographs of the 30.5-centimeter (12-inch) Hardcore FRP material. The first micrograph (A) on the left is the material in the original condition it was received. The center micrograph (B) is a sample at moisture saturation, and the sample shown on the right (C) is after freeze-thaw cycling. All micrographs are magnified times 1000. In the condition received (A), the micrograph consists of circular components with entire margins and some variation in size. Also shown are longitudinally oriented fibers. At moisture saturation (B), the micrograph consists of circular components, many with evidence of fraying at the margins. After the freeze-thaw cycles (C), the micrograph is similar to that of the original and consists of densely packed circular components of varying size. No evidence of damage is visible.

Figure 75. Graph. Estimated long-term axial capacity of the 12-inch Lancaster pile. This figure presents the axial load-strain response curve for a 30.5-centimeter (12-inch)Lancaster pile. The X-axis is axial strain in units of microstrain (0 to 16) and the Y-axis is axial load in kilonewtons (0 to 5000). Two curves are shown originating from zero. A solid line represents a typical axial load-strain response curve determined from the Fam and Rizkalla model and based on the hoop properties of the original 30.5-centimeter (12-inch) Lancaster FRP material as it was received. Between an axial strain of 0 and 1 microstrains, the axial load increases rapidly to 1500 kilonewtons. Thereafter, the slope of the line is less steep. At approximately 14 microstrains, the line reaches 4500 kilonewtons. Paralleling this load-strain response curve is a second dotted-line long-term load-strain curve that represents a 5-percent reduction in axial structural capacity. This second curve suggests that the impact of degradation of the FRP mechanical properties on the long-term axial structural capacity of the concrete-filled FRP composite pile is small due to the fact that the majority of the capacity contribution is from the concrete infill. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 76. Graph. Estimated long-term flexural capacity of the 12-inch Lancaster pile. This figure presents the moment-curvature response for a 30.5-centimeter (12-inch)Lancaster pile. The X-axis is curvature in units of the product of 1 over meters times 10 to the minus 3 (0 to 140) and the Y-axis is moment in kilonewton-meters (0 to 250). Two curves are shown originating from zero. A solid line represents the short-term moment-curvature response determined from the Fam and Rizkalla model and based on the longitudinal properties of the original 30.5-centimeter (12-inch) Lancaster FRP material as it was received. The moment-curvature relationship is somewhat linear, reaching a moment of 200 kilonewton-meters at a curvature of about 110. A second similar moment-curvature response, representing a 24-percent reduction in long-term structural flexural capacity, is shown below the short-term response solid line as a dashed line. This second curve suggests that the impact of degradation of the FRP mechanical properties is significant for the flexural capacity because the FRP shell contributes most of the capacity on the tension side of the pile in flexion. The conversion factor for kilonewton-meters is 1 kilonewton meter equals 737.57 poundforce-feet.

Figure 77. Map. Location map of the Route 40 Bridge project in Sussex County, Virginia (Fam, et al., 2003). This figure consists of two small maps. On the left is a map of the State of Virginia showing several interstate highways and the general location of the bridge project in Sussex County in the southeastern part of the State. On the right is an area map of Sussex County showing the Nottoway River and its tributaries, Interstate 95, State Route 40, State Route 657, the Town of Stony Creek near the junction of Routes 657 and 40, and the location of the Route 40 Bridge.

Figure 78. Photo. Former Route 40 Bridge. The figure is a color photograph of the old Route 40 Bridge across the Nottoway River. The photograph shows the steel truss that is supported by concrete piers.

Figure 79. Photos. Signs of deterioration of the former Route 40 Bridge (Fam, et al., 2003). This figure consists of three color photographs demonstrating cracks and deterioration in the supports. The first photograph (A) is a concrete column of the bridge pier, revealing full-height vertical cracks. Photograph (B) is a closeup of a concrete bearing seat, revealing primarily vertical cracks and spalling. The third photograph (C) shows map cracking horizontally along an abutment bearing seat.

Figure 80. Illustration. Schematic of the new Route 40 Bridge. This figure consists of two simple line drawings illustrating the schematic plans of the new bridge in elevation view and in plan view. The first illustration is an elevation view of the bridge, identifying its overall length of 85.34 meters (280 feet), and a 5-span concrete slab supported by 4 piers and 2 end abutments. The illustration also shows the positions of the battered piles alongside the reinforced concrete beam-type pile structure. The reinforced concrete pile cap is shown joining the piers to the concrete slab. The plan view shows the four piers and the spaced distance between them. From left to right, pier number 1 is located 11.25 meters (37 feet) from the left end abutment. Pier number 2 is located 18.52 meters (61 feet) from pier number 1. Pier number 3 is located 18.52 meters (61 feet) from pier number 2. Pier number 4 is located 18.52 meters (61 feet) from pier number 3 and 18.55 meters (61 feet) from the right end abutment. The plan view also shows the structural components of each pier. Piers number 1, 3, and 4 each consist of 10 prestressed solid square-column concrete piles. Pier number 2 consists of 10 concrete-filled FRP circular tubes for support. The test pile area is shown between piers number 1 and 2.

Figure 81. Photo and Illustration. Concrete-filled tubular piles. This figure consists of two color photographs of the concrete-filled FRP tubes (A) and a simple line drawing of the laminate structure of the composite piles (B). The first photograph depicts four of the tubular composite piles as they are being installed at pier number 2 of the new bridge. The second photograph is a view of the tubular composite piles from below the pier structure, showing 6 of the 10 supporting piles. The line drawing (B) below the photographs illustrates the laminate structure of the composite pile. The composite consists of 5 FRP layers, each 1.13 millimeters (0.044 inches) thick and together comprising the structural wall thickness of the FRP tube. From left to right, in the longitudinal direction, the angles of the fibers in each layer, top to bottom, are identified as positive 35 degrees, negative 35 degrees, positive 85 degrees, positive 35 degrees, and negative 35 degrees. Below these five layers is a 1.68-millimeter- (0.066-inch-) thick inner liner.

Figure 82. Graph. Stress-strain response of concrete used in the composite pile. This figure is a stress-strain graph for the concrete mix used to fill the FRP tube. The X-axis consists of tension strain and compression strain, both in meters per meter. The tension strain is shown between negative 0.002 and 0 meters per meter. Compression strain is shown between 0 and 0.005 meters per meter. The Y-axis is stress in megapascals (negative 10 to positive 70). Data tracings displayed on the chart are from the concrete model and two experimental compression tests on concrete core samples from the composite test pile. The three tracings rise from 0 megapascals at 0.000 meters per meter to approximately 60 megapascals at a compression strain of approximately 0.004 meters per meter. Just to the left of 0.000 meters per meter, the concrete model tracing dips to approximately minus 4 megapascals and then rising abruptly to just below 0 megapascals. At this point, the tracing is labeled "residual tensile strength due to tension stiffening." The tensile strength of the concrete mix, based on Brazilian tensions tests, is 4.35 megapascals.

Figure 83. Illustration. Reinforcement details of prestressed concrete pile. This figure consists of two simple line drawings showing details of the prestressed concrete pile. The first drawing is a cross section of the 508-millimeter- (20-inch-) square concrete pile showing the 14 seven-wire strands arranged evenly around and 75 millimeters (2.9 inches) in from the perimeter of the pile. Each seven-wire strand is approximately 13 millimeters (0.5 inches) in diameter. The second figure is a horizontal view of the concrete pile showing the overall length of 13.1 meters (43 feet), and the locations of the wire spiral turns along the length of the pile. The wire ties are identified as number 16M (number 5) gage spiral wire. At both ends of the column, the wire is shown with approximately 5 turns in 25 millimeters (0.98 inches). The next segments toward the middle show approximately 16 turns in 75 millimeters (2.9 inches).

Figure 84. Illustration. Simplified soil stratigraphy near test pile area. This figure consists of an illustrated soil column, identifying general soil layers below the surface, and an adjacent chart displaying standard penetration test blow counts versus depth. The soil column depicts five defined layers that correspond to the adjacent chart of blow counts. From top to bottom, they are (1) between the surface and 3.5 meters (11.5 feet) depth, a brown medium-fine loose sand with some fine gravel and traces of silt; the blow count average is about 5 per 0.3 meters (about 1 foot); (2) between about 3.5 meters (11.5 feet) and 6 meters (19.7 feet), a brown fine-grained medium-dense sand, silty with some gravel; the blow count is about 16 per 0.3 meters (1 foot); (3) between 6 meters (19.7 feet) and 8 meters (26.2 feet), a gray fine-grained medium-dense sand, silty with some gravel; the blow count average is about 21 per 0.3 meters (1 foot); (4) between 8 meters (26.2 feet) and about 10 meters (33 feet), a green very stiff to hard clay, with a trace of silt; the blow count is about 40 per 0.3 meters (1 foot); and (5) from about 10 meters (33 feet) to the lower limit of 16 meters (52 feet), a green hard silty clay layer; the blow count varies between 60 and 88.

Figure 85. Photo. Fabrication of prestressed concrete pile. This figure is a color photograph of a standard prestressed concrete pile being fabricated at the precasting plant. It shows the pile form, wire strands and spiral ties of the pile, and construction workers pouring concrete into the pile form.

Figure 86. Photos. Fabrication of concrete-filled FRP piles. This figure consists of three color photographs illustrating the fabrication of concrete-filled FRP piles. The top photograph (A) shows an FRP tube, supported by a steel beam along the entire length of the tube, held in an inclined position while the tube is being filled with the concrete mix. Photograph B shows a different view of the same scene shown in photograph A. Seven tubes are lying on the ground. The end of each is sealed with a wooden plug held in place by three metal straps attached to the tube. Photograph C shows an eight-point lifting device lifting a filled tube.

Figure 87. Photos. Driving of test piles. This figure consists of two color photographs depicting the composite pile and the prestressed pile being driven into the ground by a hydraulic impact hammer. Photograph A on the left shows the tubular composite pile vertically attached to a crane-supported hydraulic hammer and partially embedded in the soil. On the right side, photograph B shows the prestressed square pile vertically attached to a hydraulic hammer and partially embedded in the soil. In the foreground of photograph B, a tubular composite pile is shown fully embedded in the ground.

Figure 88. Graph. Driving records for test piles. This figure compares the driving records of the composite pile and the prestressed concrete pile. The X-axis is number of blows per 0.25 meters (9.8 inches). The Y-axis is the depth in meters below the driving template. At the end of driving, approximately 10.5 meters (34.4 feet) below the driving template, the composite pile required 60 blows per 0.25 meters (9.8 inches). The prestressed pile required about 32 blows per 0.25 meters (9.8 inches) to reach approximately the same depth. The chart indicates that, overall, compared to the prestressed pile, more blows are required to drive the composite pile.

Figure 89. Graphs. End-of-driving PDA recordings. This figure consists of two graphs presenting the end-of-initial-driving records from the pile driving analyzer (PDA) for the composite pile (A) and for the prestressed pile (B). The X-axis on both charts is time in milliseconds (0 to 51.2). The left Y-axis is force in kips (to 3000) and the right Y-axis is velocity in feet per second (0 to 16.6 for the composite pile and 0 to 17.8 for the prestressed pile). Graphs A and B suggest that both types of piles have similar dynamic behavior. Each graph reveals two distinct force peaks and a velocity peak.

Figure 90. Illustration. Test pile instrumentation. This figure shows the locations of instrumentation fitted to the prestressed and the composite test piles. It consists of a vertically oriented generic test pile and two cross sections representing each type of test pile. The vertical pile shows the locations of six strain gages oriented in the axial direction at three depths along the 13.1-meter- (43-foot-) long pile. Three pair of strain gages, each set positioned at opposite sides of the pile, are located at 3.87 meters (12.7 feet), 6.95 meters (22.8 feet), and 11.58 meters (38 feet) down the length of the pile. A string of eight lateral motion sensors is shown running through the center of the test pile. The upper six sensors are located 1 meter (3.3 feet) apart and the lower two are located 2 meters (6.56 feet) apart. An embedded accelerometer is shown installed 0.3 meters (0.98 feet) above the bottom of the test pile. A second illustration is a circular cross-sectional view of the composite pile, shown with the lateral motion sensor located in the center, and the pair of axially oriented strain gages, each positioned 150 millimeters (5.9 inches) in from the exterior of the pile. The final illustration is a square cross-sectional view of the prestressed pile showing the location of the lateral motion sensor in the center, and a pair of axially oriented strain gages, each positioned 75 millimeters (2.9 inches) in from the exterior of the pile.

Figure 91. Photo. Axial load test setup using the Statnamic device. This figure is a color photograph of the Statnamic Testing System used to conduct axial load tests on the composite and prestressed test piles in the field. The photograph shows the reaction mass; the pressure chamber directly below the reaction mass; the load cell below the pressure chamber; and the test composite pile, directly below the load cell. A laser-beam device, mounted in position to measure the test pile head displacement, is also shown in the photograph.

Figure 92. Graph. Pile head displacement versus equivalent axial static load. This graph depicts the equivalent static load versus the pile head axial displacement for the three axial load testing cycles conducted on the prestressed and composite test piles. The X-axis is equivalent static load in kilonewtons (0 to 5000) and the Y-axis is pile head axial displacement measured in millimeters (0 to negative 20) (0 to negative 0.79 inches). The data tracing for the composite pile shows an ultimate load capacity of about 4300 kilonewtons at a head displacement of about 18 millimeters (0.71 inches). A maximum load capacity of about 4190 kilonewtons applied to the prestressed pile corresponds to a pile head displacement of about 13 millimeters (0.51 inches). For both test piles, a rapid increase in the pile displacement near the end of the third cycle axial load testing suggests that the geotechnical capacity of the piles was fully mobilized.

Figure 93. Graph. Axial load-axial strain behavior of test piles. This graph presents the relationship between axial static load and axial strain for each of the two test piles. The X-axis is axial strain times 10 to the minus six (0 to 800). The Y-axis is equivalent axial static load in units of kilonewtons (0 to 5000). Data for both the composite pile and the prestressed pile reveal a linear relationship between axial strain and axial load. At a maximum load of about 4300 kilonewtons, the axial strain of the composite round-column pile is around 600 times 10 to the minus six. For the prestressed square-column pile, a maximum axial load of about 4200 kilonewtons results in an axial strain of about 675 times 10 to the minus six. The behavior of both piles is generally similar, but the composite pile exhibits more stiffness than the prestressed pile. The chart also shows the design load as a broken horizontal line at 667 kilonewtons.

Figure 94. Illustration and Graph. Variation of axial strain along pile length for three Statnamic loads. This figure shows the locations of the axial strain gages along the length of a test pile and includes a graph relating these gages to their strain measurements for three test cycles. The X-axis is axial strain times 10 to the minus six (0 to 800) and the Y-axis is distance from the tip of the pile (0 to 12 meters) (0 to 39 feet). The variations in axial strain along the length of the composite and prestressed test piles are shown for three test cycles. The measured strain along the length of the pile is similar for both the composite and prestressed materials, and the strain for both types decreases with increasing depth of the pile.

Figure 95. Photos. Lateral Statnamic setup at the Route 40 project. This figure consists of two color photographs of the Statnamic Testing System used to conduct lateral load tests on the composite and prestressed test piles in the field. The first photograph shows the horizontally mounted system in place at the bridge site. It shows the pressure chamber between the composite pile and the reaction mass. The second photograph shows the reaction mass being connected to the pressure chamber.

Figure 96. Graphs. Displacement time histories at the loading point for both piles (Pando, et al., 2004). This figure consists of two graphs presenting the displacement measurements versus time for the prestressed pile (A) and the composite pile (B). The X-axis in both charts is time in seconds (0 to 1) and the Y-axis is displacement in meters (negative 0.2 to positive 0.2) (negative 0.66 feet to positive 0.66 feet). Data tracings are included for the four load cycles applied to each of the test piles. For both test piles, the graphs indicate that the period of oscillation increases with increasing displacement amplitude. Compared to the data tracing for the prestressed pile shown in chart A, the larger displacement amplitude and longer period shown in chart B suggest a less stiff lateral response for the composite pile.

Figure 97. Graphs. Peak lateral displacement profiles for both test piles at different cycles of Statnamic load. This figure consists of two graphs presenting peak lateral displacement profiles for the four Statnamic load test cycles applied to the two test piles. The X-axis is peak lateral deflection in meters (0 to 0.4) (0 to 1.3 feet) and the Y-axis is depth below the load point in meters (positive 1 to negative 10) (positive 3.3 to negative 33 feet). The variations in peak lateral deflection are shown for each of the four load cycles and for each type of pile. The data in these two charts indicate that both test piles appear to form a hinge at about 5 meters (16 feet) below the point of loading. Maximum lateral deflection shown for the prestressed concrete pile is 0.3 meters (0.98 feet) and for the composite pile it is 0.35 meters (1.15 feet). For both types of piles, the deflection profile along the depth of the piles is nearly bilinear for load cycles 3 and 4, with a change in slope detected at a depth of about 4.8 meters (16 feet) below the loading point.

Figure 98. Graph. Calculated static and dynamic (static plus damping) resistances for both test piles. This figure presents the peak static soil resistance and the peak dynamic soil resistance calculated for the each of the four lateral load cycles applied to the prestressed pile and to the composite pile. The X-axis is displacement in meters (0 to 0.4) (0 to 1.3 feet) and the Y-axis is load in kilonewtons (0 to 200) (0 to 45,000 poundforce). For both types of piles, a comparison of the static soil resistance with the total (static plus damping) soil resistance indicates that damping accounts for only a small amount of the total resistance. The graph also indicates that, at a lateral load of about 50 kilonewtons (11,250 poundforce), there is a significant change in the slope of the load versus the deflection curve for the composite pile. For the prestressed pile, the change in slope of the load versus the deflection curve occurs at a lateral load of about 125 kilonewtons (28,125 poundforce).

Figure 99. Graph. Moment-curvature responses for composite and prestressed concrete piles. This figure presents the calculated and experimentally derived moment-curvature responses for the prestressed and composite test piles. The X-axis is curvature in radians per kilometer (0 to 40) and the Y-axis is moment measured in kilonewton-meters (0 to 700). For a study performed by Fam (2000), the chart reveals that there is good agreement between the calculated and experimental data. The moment-curvature data, calculated for the composite pile used in this study, are displayed as mostly linear and the response is stiff until a moment of about 110 kilonewton-meters is reached. Beyond this value, the composite pile exhibits less stiffness. The moment curvature response for the prestressed pile, derived from modeling predictions, is somewhat logarithmic. The curve achieves a nearly flat moment of 500 kilonewton-meters at a curvature of approximately 15 radians per kilometer.

Figure 100. Graph. Computed and measured lateral load-displacement response for both test piles (Pando, et al., 2004). This figure compares the predicted and measured lateral load-displacement responses for the prestressed and composite piles. The X-axis is displacement in meters (0 to 0.4) (0 to 1.3 feet) and the Y-axis is applied load in kilonewtons (0 to 200) (0 to 45,000 poundforce). The data show that the predicted and measured responses for the composite pile track very well, with a lateral load of about 120 kilonewtons (27,000 poundforce) resulting in a displacement of 0.3 meters (0.98 feet). For the prestressed pile, the predicted and measured responses agree well initially, although the calculated minimum load is less than the measured maximum load, but after a load of about 100 kilonewtons (22,500 poundforce), the measured response is significantly higher than the predicted response.

Figure 101. Illustration. Details of pile head showing the bars used to connect the pile to cap beam. This figure consists of two sets of drawings, each representing the pile head and a cross section of the top end of each of the two types of test piles, the square-column prestressed pile and the circular-tube composite pile. The cross-sectional views illustrate the positions of the eight 25.4-millimeter- (1-inch-) diameter holes drilled parallel to the longitudinal axis and to a depth of 460 millimeters (18 inches). In the prestressed pile, the 8 holes are shown equally spaced along a 330-by-330-millimeter- (13-by-13-inch-) square perimeter within the 508-by-508-millimeter (20-by-20-inch) square top of the pile. In the composite pile, the 8 drilled holes are shown equally spaced along a 447-millimeter- (17.6-inch-) diameter circle within the 625-millimeter- (24.6-inch-) diameter composite pile. The longitudinal views show the positions of the 1219-millimeter- (48-inch-) long Number-7 steel rebars inserted into the eight holes. In both the prestressed and composite piles, the rebars are shown embedded in the pile to a depth of 457 millimeters (18 inches).

Figure 102. Illustration. Connection of composite piles to cap beam at Pier Number 2. This figure consists of two drawings depicting (A) the elevation of Pier Number 2, and (B) the longitudinal section of the pier at a location where the composite pile is connected to the cap beam. The first illustration (A) shows Pier Number 2, consisting of the horizontal cap beam supported by 10 vertical composite piles. The drawing also indicates the location enlarged in the longitudinal section shown in (B). The next drawing (B) is a longitudinal section through a point in the cap beam where a composite pile is attached. This illustration shows the Number-7 steel rebars embedded in the cap beam. The section of the cap beam shown is 1 meter (39.4 inches) wide and 1 meter (39.4 inches) high. Drawing (B) also shows rebars throughout the cap beam.

Figure 103. Photos. Pier Number 2 including the composite piles and reinforced concrete cap beam. This figure consists of two color photographs showing the reinforced concrete cap beam of Pier Number 2 and a closeup view of the cap beam joined to a composite pile.

Figure 104. Photo. The new Route 40 Bridge over the Nottoway River in Virginia. This figure is a color photograph of the new Route 40 Bridge, showing 8 of the 10 composite piles of Pier Number 2. Two other bridge piers, constructed of the square-column prestressed concrete piles, are also visible in the photograph.

Figure 105. Map. Location map of the Route 351 Bridge in Hampton, Virginia. This figure consists of two small maps. On the left is a map of the State of Virginia showing the major interstate highways and the general location of the bridge project in Hampton, located in the southeastern part of the State. On the right is an enlarged view of the project area showing the Hampton Roads Bridge-Tunnel, and Interstate 64 and State Route 351 where they cross the Hampton River.

Figure 106. Photo. Aerial view of the Route 351 Bridge in Hampton, Virginia. This figure is a Globexplorer color aerial photograph of the original Route 351 Bridge over the Hampton River. The photograph shows the bridge located beneath and at an angle (southwest to northeast) opposite to the Interstate-64 overpass (northwest to southeast).

Figure 107. Photos. Wide-angle views of the original Route 351 Bridge. This figure consists of two color wide-angle photographs of the original Route 351 Bridge, taken from the north side. Both photographs show most of the length of the bridge and its position beneath the Interstate-64 overpass. The first photograph also clearly shows the piers being supported by three piles. The second photograph is a closer view of the bridge showing the same features.

Figure 108. Photo. Signs of deterioration of the original Route 351 Bridge. This figure is a color photograph of the Route 351 Bridge, taken from the south side, showing deterioration of the pier pilings and also in the superstructure. Dotted lines show the areas of deterioration.

Figure 109. Illustration. Schematic of the new Route 351 Bridge. This figure consists of two simple line drawings illustrating the new Route 351 Bridge in elevation view and in plan view. The first illustration is an elevation view of the proposed bridge, showing the superstructure supported by 12 piers and 2 end abutments. The illustration shows the first four spans on the west side of the bridge supported by steel plate girder spans and the remaining bridge deck supported by prestressed concrete beams. The plan view shows the 12 piers, the 2 end abutments, their orientation, and the distance between each pier. Each pier, except number 2, is shown supported by seven prestressed concrete piles. The orientation of each of the 1-to-4 battered piles is also indicated by arrows on the pier schematics. From west to east, pier number 1 is located 15.2 meters (49.9 feet) from the west end abutment. Pier number 2 is located 26.6 meters (87.3 feet) from pier number 1. Pier number 3 is located 25.4 meters (83.3 feet) from pier number 2. Pier number 4 is located 21.29 meters (69.8 feet) from pier number 3. Piers 5 through 12 are each located 15.49 meters (50.8 feet) apart and pier 12 is also 15.49 meters (50.8 feet) from the right end abutment. The plan view also shows the location of an instrumented prestressed concrete pile in the center of pier number 10 and an instrumented FRP composite pile in the center of pier number 11.

Figure 110. Illustration. Test pile cross section details. This figure consists of three drawings comparing cross sections of the prestressed concrete pile (A), the FRP pile (B), and the plastic pile (C). The cross section of the prestressed concrete pile (A) illustrates the arrangement of the 12.7-millimeter- (0.5-inch-) diameter steel strands. The 16 steel wire strands are shown equally spaced along a 432-millimeter- (16.8-inch-) square perimeter within the 610-millimeter- (24-inch-) square pile column. Text below the illustration states that the strands are tied to a Number 15M-gage-wire external spiral with a 0.15-meter (5.9-inch) pitch. The cross section of the 622-millimeter- (24.5-inch-) diameter composite pile (B) shows the 10.7-millimeter- (0.42-inch-) thick outer FRP shell and the inner concrete core reinforced with steel rebar. The reinforcement consists of 14 Number 25M steel bars equally spaced along a 506-millimeter- (19.7-inch-) diameter circle within the concrete core. Text below the illustration states that the rebar is tied to a Number 9M-gage-wire external spiral with a 0.15-meter (5.9-inch) pitch. The cross section of the 592-millimeter- (23.3-inch-) diameter plastic pile (C) illustrates the recycled MDPE plastic matrix reinforced with 24 Number 25M steel bars. The bars are shown equally spaced along a 506-millimeter- (19.7-inch-) diameter circle within the plastic matrix. Text below the illustration states that the rebar is welded to a Number 9M-gage-wire internal spiral with a 0.23-meter (9-inch) pitch.

Figure 111. Graphs. Test pile material properties. This figure consists of seven graphs summarizing material properties of the three types of test piles. Part A of this figure includes two stress-strain charts for the prestressed concrete test pile. The X-axis of the chart for the concrete is strain in meters per meter (0 to 0.003) (0 to 0.010 feet) and the Y-axis is compressive stress in megapascals (0 to 60). The concrete stress-strain diagram shows two results of nearly linear relationships to a stress of about 50 megapascals at a strain of about 0.00225 meters per meter (0.074 feet), at which point there are indications of leveling off. For the steel component of the prestressed concrete pile, the X-axis is strain in meters per meter (0 to 0.05) (0 to 0.16 feet) and the Y-axis is tensile stress in megapascals (0 to 2,000). The curve for the prestressed steel strand consists of two components: a linear relationship to a tensile stress of about 1,750 megapascals at a strain of 0.01 meters per meter (0.033 feet); and a horizontal component at a tensile stress of about 1,800 megapascals between a strain of 0.01 and 0.04 meters per meter (0.033 and 0.13 feet). Text associated with this chart states that the diagram represents a 7-wire strand, 12.7 millimeters (0.5 inches) in diameter, with a low relaxation and a yield of 1,850 megapascals. Part B of this figure includes two stress-strain charts for the FRP test pile. The X-axis of the chart for the concrete fill is strain in meters per meter (0 to 0.002) (0 to 0.0066 feet) and the Y-axis is compressive stress in megapascals (0 to 40). The diagram shows a nearly linear relationship to a stress of about 35 megapascals at a strain of about 0.0014 meters per meter (0.0046 feet). For the FRP shell, the stress-strain relationship is compared to that for Grade 420 steel. The X-axis for this chart is strain in meters per meter (0 to 0.02) (0 to 0.066 feet) and the Y-axis is tensile stress in megapascals (0 to 500). The diagram for the FRP shell is a nearly perfect linear relationship between tensile stress and strain. The diagram for Grade 420 steel, included for comparison, is again a two-part curve, similar to that of the steel strands in the prestressed concrete pile (A). Part C of this figure consists of three charts showing two stress-strain diagrams for recycled plastic and one for Grade 420 steel. The X-axis of the two charts for plastic is strain in meters per meter (0 to 0.10) (0 to 0.33 feet) and the Y-axis is compressive strength in megapascals (0 to 12). The relationships in both charts appear to be somewhat logarithmic in shape. At a stress between 9 and 10 megapascals, the strain is between 0.08 and 0.09 meters per meter (0.26 and 0.29 feet). The final chart in part C shows the stress-strain relationship for the Grade 420 steel reinforcements). The X-axis is strain in meters per meter (0 to 0.01) (0 to 0.33 feet) and the Y-axis is tensile stress in megapascals (0 to 500). The diagram consists of a two-part curve and is very similar to the stress-strain relationship shown for Grade 420 steel in part B of this figure. The conversion factor for megapascals is 1 megapascal equals 145 poundforce per square inch.

Figure 112. Graph. Axial load-axial strain behavior of test piles. This figure presents the axial load-axial strain relationships for the three types of test piles. The X-axis is axial strain in meters per meter (0 to 0.002) (0 to 0.0066 feet) and the Y-axis shows axial load in units of kilonewtons (0 to 18,000). All three piles exhibit linear strain-load relationships. The line for the plastic pile has the smallest slope, the line for the prestressed concrete has the largest slope, and the line for the FRP composite pile is just below that of the concrete pile. The prestressed concrete pile exhibits the highest axial stiffness (EA) at 8.2 times 10 to the 6 kilonewtons. The lowest axial stiffness, 3.2 times 10 to the 6 kilonewtons, is associated with the plastic pile. The axial stiffness of the FRP composite pile is 7.36 times 10 to the 6 kilonewtons. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 113. Graphs. Flexural characteristics for the three test piles. This figure consists of two graphs presenting the moment-curvature response (A) and the flexural stiffness as a function of applied moment (B) for the prestressed concrete pile, the FRP composite pile, and the plastic pile. The X-axis of the first chart (A) is curvature in radians per kilometer (0 to 30) and the Y-axis is moment measured in kilonewton-meters (0 to 1,200). The shape of the moment-curvature response for the prestressed concrete pile appears logarithmic, with a maximum moment at about 700 kilonewton-meters between 5 and 25 radians per kilometer. For the plastic pile, a two-part curve is shown. The initial part is linear to a curvature of about 10 radians per kilometer at a moment of about 625 kilonewton-meters. Between a curvature of about 10 and 30 radians per kilometer, the moment-curvature response for the plastic pile then levels out between 625 and about 800 kilonewton-meters. The response for the FRP composite pile appears similar to the response for the plastic pile but, rather than leveling out, continues a gradual increase to about 1100 kilonewton-meters at 25 radians per kilometer. Part B of this figure presents flexural stiffness as a function of applied moment. The X-axis is moment in kilonewton-meters (0 to 1,200) and the Y-axis is flexural stiffness (EI) in kilonewton-meters squared (0 to 400,000). A two-part curve is shown for the prestressed concrete pile. Between a moment of 0 and about 350 kilonewton-meters, the response curve for the prestressed concrete pile is essentially flat at a flexural stiffness of about 36,000 kilonewton-meters squared. Between about 350 and 725 kilonewton-meters, the curve decreases linearly to about 4,000 kilonewton-meters squared. For the FRP composite pile, the flexural stiffness curve is flat at about 185,000 kilonewton-meters squared between a moment of 0 and about 200 kilonewton-meters. It then declines linearly to a moment of 250 kilonewton-meters and a flexural stiffness of about 140,000 kilonewton-meters squared. Thereafter, the curve declines very gradually to a moment of around 1,150 kilonewton-meters and a flexural stiffness of 5,000 kilonewton-meters squared. The flexural stiffness curve for the plastic pile is essentially flat at about 7,500 kilonewton-meters squared between a moment of 0 and about 700 kilonewton-meters. Between 700 and 800 kilonewton-meters, the flexural stiffness declines to about 2,000 kilonewton-meters squared. The conversion factors are: 1 kilonewton-meter equals 737.57 poundforce-feet; 1 radian per kilometer equals 0.621 radian per mile; and 1 kilonewton-meter squared equals 2,421 poundforce-feet squared.

Figure 114. Map. Location of test pile site at the Route 351 Bridge. This figure is a color map showing the location of the test pile area in the vicinity of the Hampton River in the City of Hampton, Virginia. The test pile area is shown on a spit of land extending into the Hampton River just south of the existing Route 351 Bridge. Also shown crossing the Hampton River are Interstate-64 and Route 351.

Figure 115. Charts. Simplified soil stratigraphy near test pile area. This figure consists of a simplified soil column showing depth and stratigraphy, and three charts showing boring and probe test results with depth. Depth is shown on the left from 0 at the top to 30 meters (98 feet) at the bottom. Between the surface and about 1 meter (3.3 feet), the soil material is described as fill and silty fine sand (SM). Between 1 meter (3.3 feet) and about 13 meters (42.6 feet), the material is primarily fine-grained medium-dense silty sand, with shell fragments, and either brown or gray in color (SM). From about 13 meters (42.6 feet) to 16 meters (52.5 feet), the soil consists of sandy stiff clay, gray in color, with traces of shell fragments (CL). Below this layer, between 16 and 19 meters (52.5 and 62.3 feet), the soil is characterized by medium dense fine-grained gray silty sand (SM). Between 19 and about 23 meters (62.3 and 75.5 feet), the soil layer is described as medium dense clayey to silty sand, with traces of shell fragments and gray in color (SM-SC). In the layer between 23 and 30 meters (75.5 and 98 feet), the material is medium-dense to dense fine-grained silty sand, gray in color, with shell fragments. Adjacent to the soil stratigraphy is a chart of standard penetration test N-values. The X-axis is the N-value (0 to 30) and the Y-axis is depth in meters (0 to 30) (0 to 98 feet), corresponding to the soil stratigraphy depths. The results for two field tests are shown in a scatter plot. The lowest values, between about 3 and 15 are shown corresponding to the upper 8 meters (26 feet). Below that depth, the field N-values range from about 10 to 30 and are somewhat scattered between 8 and 15 meters (26 and 49 feet). At the lower depths, from about 15 to 30 meters (49 to 98 feet), the field N-values range more narrowly between about 20 and 30, with very good agreement between the two tests. Adjacent to the SPT chart is a depth profile of the tip resistance and another profile showing the sleeve resistance, both obtained from cone penetrometer tests. The X-axis of the tip resistance chart is bars (0 to 400) and the X-axis of the sleeve friction profile is also bars (0 to 4). The tip resistance data include the results of four tests, all in very good agreement. The tip resistance appears to be greatest between 23 and 30 meters (75.5 and 98 feet). The results of the four sleeve friction tests are more inconsistent and do not show an obvious trend.

Figure 116. Illustration. Pile load test layout. This figure is a diagram of the layout of the borings and probe used to conduct load tests on the three test piles. The scale shown is 4 centimeters (1.6 inches) equal 5 meters (16.4 feet). The test piles, boring locations, and probe locations are shown in vertical alignment. The arrangement of the three test piles along this vertical axis, from top (north) to bottom (south), consists of the hardcore composite pile, the PPI recycled plastic pile, and the prestressed concrete pile. Directly north of and adjacent to the hardcore composite test pile is the location of a cone penetrometer test probe (CPT034). On the same vertical axis and about 5 meters (16.4 feet) north of the hardcore composite test pile is the location of a standard penetration test boring (SPT-1). The SPT-1 boring site is approximately in the center of a quadrangle formed by four HP 14-by-89 reaction piles. The reaction pile in the southwest corner is located approximately 2.85 meters (9.35 feet) northeast of the hardcore composite pile. Approximately 3 meters (9.8 feet) south of the hardcore composite test pile is another cone penetrometer test probe (CPT-1) location, also in the approximate center of four HP 14-by-89 reaction piles. On the vertical axis about 8 meters (26 feet) south of the hardcore composite test pile is the location of the PPI recycled plastic test pile. Four meters (13 feet) below the plastic pile is the location of another standard penetration test boring (SPT-2), which is located in the approximate center of four HP 14-by-89 reaction piles. Four meters (13 feet) south of the SPT-2 boring and 2.85 meters (9.35 feet) southeast of the southwest corner reaction pile is the location of the prestressed concrete test pile. Immediately below and adjacent to the concrete test pile is the location of a cone penetrometer test probe (CPT033). A DMT probe (DMT-1) is located about 3 meters (9.8 feet) on the vertical axis south of the prestressed concrete test pile. Another cone penetrometer test probe (CPT-2) is approximately 1 meter (3.3 feet) south of the DMT probe location. The DMT-1 and the CPT-2 are approximately centered in a quadrangle of four HP 14-by-89 reaction piles.

Figure 117. Illustration. Instrumentation layout for prestressed concrete test pile. This figure consists of two illustrations depicting the vertical layout and a cross section of the instrument-fitted prestressed concrete test pile. The vertical diagram shows the locations of 16 "sister-bar" strain gages along the 18-meter- (59-foot-) long test pile. A pair of "sister bars," on opposite sides of the pile, is located at 6 intervals along the pile. Two additional gages are shown at the top and also at the bottom of the test pile. Four sister bars (SB-1) are located at a depth of 1.12 meters (3.7 feet) from the top of the pile. Two "sister bars" (SB-2) are shown 3.25 meters (10.5 feet) below SB-1. Two "sister bars" (SB-3) are shown at a depth of 3.05 meters (10 feet) below SB-2. A pair of "sister bars" (SB-4) is located 2.97 meters (9.7 feet) below SB-3. Three meters (9.8 feet) below SB-4 is another pair of "sister bars" (SB-5). The last set of four "sister bars" (SB-6) is located 4.01 meters (13 feet) below SB-5. The vertical diagram of the instrumented test pile also shows a 17.06-meter- (56-foot-) long inclinometer casing through the center of the pile. The cross section diagram of the instrumented test pile shows the inclinometer casing through the center of the square-column pile, and a pair of "sister bars," each mounted 95 millimeters (3.74 inches) in from the outer wall and on opposite sides of the pile.

Figure 118. Illustration. Instrumentation layout for FRP composite test pile. This figure consists of three illustrations depicting two vertical layouts and a cross section of the instrument-fitted FRP composite test pile. One vertical diagram shows the locations of instruments embedded in the concrete infill and a second vertical diagram presents the locations of the instruments embedded in or bonded to the outer FRP tube. Embedded in the concrete are 16 "sister-bar" strain gages at six intervals along the 18.3-meter- (60-foot-) long test pile. One pair of "sister bars," on opposite sides of the pile, is located at six intervals along the pile. Two additional gages are shown at the top and also at the bottom of the test pile. Four sister bars (SB-1) are located at a depth of 1.22 meters (4 feet) from the top of the composite pile. Two "sister bars" (SB-2) are shown 3.51 meters (11.5 feet) below SB-1. Two "sister bars" (SB-3) are shown at a depth of 4.32 meters (14.2 feet) below SB-2. A pair of "sister bars" (SB-4) is located 3.15 meters (10.3 feet) below SB-3. About 2.9 meters (9.5 feet) below SB-4 is another pair of "sister bars" (SB-5). The last set of four "sister bars" (SB-6) is located 2.44 meters (8 feet) below SB-5. The vertical diagram of the concrete embedded instrumented test pile also shows a 16.76-meter- (60-foot-) long inclinometer casing through the center of the pile. A second diagram of the vertical pile illustrates the locations of instruments embedded in or bonded to the FRP shell. This diagram reveals the locations of 18 foil strain gages, mounted on opposite sides of the pile and bonded to the internal wall of the FRP tube, and 10 fiber optic sensors, mounted on opposite sides of the pile and embedded in the FRP tube. The two fiber optic cables, on opposite sides of the pile, are shown exiting the pile at 1.14 meters (3.74 feet) from the top of the pile. The first pair of foil strain gages (IG-1) is located at a depth of 0.81 meters (2.7 feet) below the fiber optic cable exits. The next set of foil gages (IG-2) is mounted 0.94 meters (3.1 feet) below IG-1 and at the same depth on the pile as the first pair of fiber optic sensors (FO-1). Approximately 1.83 meters (6 feet) below IG-2 is the location of another set of foil gages (IG-3). A pair of fiber optic sensors (FO-2) is mounted at 1.68 meters (5.5 feet) below the location of IG-3. A set of foil strain gages (IG-4) is located 1.78 meters (5.8 feet) below the location of the FO-2 soil gages. A pair of fiber optic sensors (FO-3) and a pair of foil strain gages (IG-5) are mounted 0.86 meters (2.8 feet) below IG-4. Approximately 1.75 meters (5.7 feet) below the location of the IG-5 gages, another pair of foil strain gages (IG-6) is bonded to the FRP tube. A set of fiber optic sensors (FO-4) is embedded in the FPR tube about 1.22 meters (4 feet) below the location of IG-6. A set of foil gages (IG-7) is attached at a depth of 2.01 meters (6.6 feet) below FO-4. Another set of foil strain gages (IG-8) and a set of fiber optic sensors (FO-5) are located at a depth of 1.07 meters (3.5 feet) below IG-7. The final set of foil gages (IG-9) is located 1.8 meters (5.9 feet) below IG-8 and 1.41 meters (4.6 feet) above the bottom of the test pile. The cross section diagram of the instrumented test pile shows the inclinometer casing through the center of the tubular pile, and a pair of "sister bars," each mounted 58 millimeters (2.28 inches) in from the FRP outer wall and on opposite sides of the pile.

Figure 119. Illustration. Instrumentation layout for plastic composite test pile. This figure consists of two illustrations depicting the vertical layout and a cross section of the instrument-fitted plastic composite test pile. The vertical diagram shows the locations of 16 "sister-bar" strain gages along the 18.3-meter- (60-foot-) long test pile. A pair of "sister bars," on opposite sides of the pile, is located at six intervals along the pile. Two additional gages are shown at the top and also at the bottom of the test pile. Four sister bars (SB-1) are located at a depth of 1.09 meters (3.6 feet) from the top of the pile. Two "sister bars" (SB-2) are shown 3.51 meters (11.5 feet) below SB-1. Two "sister bars" (SB-3) are shown at a depth of 3.01 meters (9.9 feet) below SB-2. A pair of "sister bars" (SB-4) is located 3.11 meters (10.2 feet) below SB-3. About 2.74 meters (9 feet) below SB-4 is another pair of "sister bars" (SB-5). The last set of four "sister bars" (SB-6) is located 4.25 meters (14 feet) below SB-5. The vertical diagram of the instrumented test pile also shows an 18-meter- (59-foot-) long inclinometer casing through the center of the pile. The cross section diagram of the instrumented test pile shows the inclinometer casing through the center of the tubular pile, and a pair of "sister bars," each mounted 65 millimeters (2.6 inches) in from the outer wall and on opposite sides of the pile. The inclinometer casing is shown inside a steel pipe with an outer diameter of 141 millimeters (5.5 inches) and a wall diameter of 6.35 millimeters (0.25 inches).

Figure 120. Photos. Fabrication of prestressed concrete test pile. This figure consists of two color photographs documenting the fabrication of the prestressed concrete test pile. The first photograph, taken at the casting yard, shows the pile partially filled with concrete. In the unfinished section, the form containing the prestressed steel bars can be seen, along with the inclinometer casing in the center and one of the "sister-bar" strain gages. The second photograph shows the finished square-column prestressed concrete test pile wrapped and ready for transport.

Figure 121. Photos. Fabrication of concrete-filled FRP piles. This figure consists of three color photographs of the concrete-filled FRP test pile. The first photograph (A) shows two FRP shells, one for the test pile and one for the production pile, on the lawn at the casting yard. The second photograph (B) is a view inside the tubular FRP shell showing the foil strain gages installed in the walls of the tube. The third photograph depicts the instrumented rebar cage on a dolly and partially inserted into the FRP shell.

Figure 122. Photos. Setup used for concrete filling of FRP composite piles. This figure consists of four color photographs documenting the steps to fill the reinforced and instrumented FRP tubes with concrete. The first photograph (A) shows the prepared FRP tube attached to a crane and being raised in a vertical position alongside a steel supporting structure. The next photograph (B) shows the two prepared FRP tubes in a vertical position and secured in the steel brace. The third photograph (C) is a closeup of the L-shaped steel angle bars securing the base of the FRP piles to a concrete slab. The last photograph (D) shows the hopper and the elephant trunk that are used to pour the concrete into the FRP tubes.

Figure 123. Photos. Concrete filling of FRP composite piles. This figure consists of a series of four color photographs documenting the steps to fill the FRP tubes with concrete. The first photograph (A) shows the 2-cubic-yard bucket being filled with concrete. Photograph (B) shows the concrete-filled bucket being hoisted by a crane into position above the vertically mounted FRP tubes. The next photograph (C) is a view from below the concrete-filled bucket, showing that operators, located in an adjacent manhold lift, control the position of the bucket over the FRP shell and the filling process. Photograph (D) is a closeup of the top of the concrete-filled FRP production pile, specifically showing the sleeves used to insert the steel dowels that connect the pile to the pile cap.

Figure 124. Photos. Rebar cage of the plastic composite test pile. This figure consists of two color photographs of the tubular rebar cage for the plastic composite test pile. The first photograph is an end view showing the steel pipe used to install the inclinometer casing in the center of the cage. The second photograph is a closeup of the side of the rebar cage showing the longitudinal steel bars and spiral ties welded to the bars.

Figure 125. Photos. Manufacturing process for the plastic composite test pile. This figure consists of five color photographs depicting steps in the manufacturing of the plastic composite test pile. The first photograph (A) shows two workers loading the instrumented rebar cage into a tubular steel injection mold. The next photograph (B) is a view of the bottom end of the steel mold loaded with the rebar cage and the center steel pipe used to install the inclinometer. Photograph (C) shows the injected steel mold being submerged horizontally into a water-cooling tank. The next photograph (D) is the manufactured pile still inside the steel mold. As shown in photograph (E), the finished product is the extruded plastic test pile.

Figure 126. Graph. Driving records for test piles. This figure presents the pile driving records for the prestressed concrete pile, the plastic pile, and the FRP composite pile. The X-axis is number of pile blows (0 to 60) per 0.25 meters (0.82 feet). The Y-axis is depth below the ground surface in meters (0 to 20) (0 to 66 feet). The shapes of the three curves are somewhat inversely proportional, indicating that more blows are required to drive the pile at greater depths. To reach a depth of about 17 meters (56 feet), about 18 blows per 0.25 meters (0.82 feet) were needed for the plastic pile. On the other hand, to reach approximately the same depth, about 50 blows per 0.25 meters (0.82 feet) were required for the prestressed concrete pile. The FRP composite pile required about 22 blows per 0.25 meters (0.82 feet) to reach a depth of about 17 meters (56 feet).

Figure 127. Photos. Installation of prestressed concrete test pile. This figure consists of three color photographs documenting the installation of the prestressed concrete test pile. The first photograph (A) shows the concrete pile below the pile driver at the bridge site. The next photograph (B) depicts the pile after the completion of driving. It shows the compressed driving cushion at the top of the pile as well as instrumentation wires exiting the pile near the top. Photograph (C) is a closeup depicting restrike of the prestressed pile.

Figure 128. Photos. Installation of FRP composite test pile. This figure consists of three color photographs documenting the installation of the FRP composite test pile. The first photograph (A) shows the FRP pile below the pile driver at the bridge site. The next photograph (B) depicts the pile after the completion of driving. It shows the top of the pile as well as instrumentation wires exiting the pile near the top. Photograph (C) also depicts the pile after completion of driving, and shows the compressed driving cushion at the top of the pile.

Figure 129. Photos. Installation of plastic composite test pile. This figure consists of three color photographs documenting the installation of the plastic composite test pile. The first photograph (A) shows the vertical plastic pile and the pile driver assembly. The next photograph (B) depicts the pile after the completion of driving. Photograph (C) is a closeup depicting restrike of the plastic pile. It also shows the protective helmet, containing a plywood cushion, in place between the pile driver and the top of the plastic pile.

Figure 130. Graphs. PDA recordings during restrike. This figure consists of three graphs (A, B, and C) presenting force and velocity monitoring data obtained from the pile driving analyzer during restrike of the prestressed concrete pile, the FRP composite pile, and the plastic composite test piles, respectively. The X-axis on each chart is time in milliseconds (0 to 50) and the Y-axis is force in kilonewtons (negative 2000 to positive 5000). Comparing the three charts, it is clear that the prestressed concrete pile and the FRP composite pile exhibit similar dynamic behavior during driving. The data indicate that the peak force, as well as the peak velocity, occurs at time zero for all three piles, although the magnitude of the peaks varies. The largest peak force, at approximately 4800 kilonewtons, was measured with the FRP composite pile. The prestressed concrete pile exhibits a peak force of about 4500 kilonewtons and the peak force recorded for the plastic pile is about 2200 kilonewtons. A velocity peak also appears on all three charts at 2L over lowercase C, representing twice the pile length (L) divided by the wave speed (lowercase C). This secondary peak is pronounced for the prestressed concrete pile and for the FRP composite pile, but less obvious for the plastic pile. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 131. Photos. PIT tests on test piles. This figure consists of three color photographs taken during pile integrity testing (PIT) conducted on the plastic composite and the FRP composite test piles. The first photograph (A) shows a technician using a hand-held hammer to apply a small-impact load on the top of the plastic pile after it has been installed. Photograph (B) depicts a technician conducting the PIT before pile installation. He is shown tapping a hand-held hammer on the end of the FRP composite pile, which is still in a horizontal position. The last photograph (C) shows the technician using a hand-held hammer to conduct the PIT on the end of the installed FRP composite pile.

Figure 132. Graph. PIT sounding on the prestressed concrete test pile before installation. This figure presents the PIT wave velocity curve for the prestressed concrete test pile before driving. The X-axis is distance in feet (0 to 75) (0 to 23 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1.0) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the chart states that the data were recorded on December 4, 2001; that the wave speed was 13,444 feet per second (4,098 meters per second); and that the exact length of the prestressed concrete pile is 59.08 feet (18 meters). The chart also includes a rectangular representation of the test pile above the X-axis, between 0 and about 60 feet (0 and 18.3 meters), corresponding to the length of the pile. The uniformity of the wave velocity data suggests that, before installation, the pile is probably free of imperfections or any damage, which would be revealed by an inflection of the wave velocity curve.

Figure 133. Graph. PIT sounding on the prestressed concrete test pile after installation. This figure presents the PIT wave velocity curve for the prestressed concrete test pile after driving. The X-axis is distance in feet (0 to 75) (0 to 23 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1.0) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the chart states that the data were recorded on December 5, 2001; that the wave speed was 13,242 feet per second (4,036 meters per second); and that the exact length of the prestressed concrete pile is 59.08 feet (18 meters). The chart also includes a rectangular representation of the test pile above the X-axis, between 0 and about 60 feet (0 and about 18.3 meters), corresponding to the length of the pile. The uniformity of the wave velocity data suggests that the pile after installation is probably free of imperfections or any damage, which would be revealed by an inflection of the wave velocity curve. Soil resistance in the pile is revealed as a toe reflection shown as a small peak on the chart.

Figure 134. Graph. PIT sounding on the FRP composite test pile before installation. This figure presents the PIT wave velocity curve for the FRP composite test pile before driving. The X-axis is distance in feet (0 to 75) (0 to 23 meters) and the Y-axis is velocity in inches per second (negative 0.2 to positive 1.0) (negative 5.1 millimeters to positive 25.4 millimeters). Text on the chart states that the data were recorded on February 20, 2002; that the wave speed was 12,959 feet per second (3,950 meters per second); and that the exact length of the FRP composite pile is 60 feet (18.3 meters). The chart also includes a rectangular representation of the test pile above the X-axis, between 0 and about 60 feet (0 and 18.3 meters), corresponding to the length of the pile. One small peak occurs in the velocity trace at about 17 feet from the top of the pile; another occurs at about 32 feet from the top of the pile. Overall, however, the uniformity of the wave velocity data suggests that, before installation, the pile is probably free of imperfections or any damage.

Figure 135. Graph. PIT sounding on the FRP composite test pile after installation. This figure presents the PIT wave velocity curve for the FRP composite test pile after driving. The X-axis is distance in feet (0 to 75) (0 to 23 meters) and the Y-axis is velocity in inches per second (negative 0.2 to positive 1.0) (negative 5.1 millimeters to positive 25.4 millimeters). Text on the chart states that the data were recorded on February 25, 2002; that the wave speed was 12,959 feet per second (3,950 meters per second); and that the exact length of the prestressed concrete pile is 60 feet (18.3 meters). The chart also includes a rectangular representation of the test pile above the data tracing, between 0 and about 60 feet (0 and about 18.3 meters), corresponding to the length of the pile. The two small peaks shown in the preinstallation PIT data (figure 134) are also visible on the post-installation tracing, suggesting that driving the pile in place did not damage the integrity of the pile.

Figure 136. Graph. PIT sounding on the plastic composite test pile before installation. This figure presents the PIT wave velocity curve for the plastic composite test pile before driving. The X-axis is distance in feet (0 to 75) (0 to 23 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1.0) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the chart states that the data were recorded on December 4, 2001; that the wave speed was 10,500 feet per second (3,200 meters per second); and that the exact length of the plastic composite pile is 60 feet (18.3 meters). The chart also includes a rectangular representation of the test pile above the X-axis, between 0 and about 60 feet (0 and 18.3 meters), corresponding to the length of the pile. Except for a negative peak at about 4 feet (1.22 m) and 2 small positive peaks at about 10 feet (3.05 meters) and 22 feet (6.7 meters), the data tracing is somewhat uniform, again suggesting a pile with few imperfections.

Figure 137. Graph. PIT sounding on the plastic composite test pile after installation. This figure presents the PIT wave velocity curve for the plastic composite test pile after driving. The X-axis is distance in feet (0 to 75) (0 to 23 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1.0) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the chart states that the data were recorded on December 5, 2001; that the wave speed was 9,620 feet per second (2,932 meters per second); and that the exact length of the prestressed concrete pile is 60 feet (18.3 meters). The chart also includes a rectangular representation of the test pile above the data tracing, between 0 and about 60 feet (0 and about 18.3 meters), corresponding to the length of the pile. Comparison with the preinstalled data shown in figure 136 shows a number of velocity peaks in the post-installation sounding data; these additional peaks may be related to potential damage caused during driving and installation of the pile.

Figure 138. Photos. Axial load test of prestressed concrete pile. This figure consists of two color photographs taken during the static axial load tests conducted on the prestressed concrete test pile. The first photograph shows the load cell and three calibrated hydraulic jacks positioned below a steel reaction beam and above the top of the installed prestressed concrete test pile. A horizontal steel reference beam is also shown touching the test pile. The second photograph is a different view of the same setup.

Figure 139. Graph. Axial load test results. This figure presents the results of the axial load tests conducted on the three types of test piles. The X-axis is axial load in kilonewtons (0 to 3,500) and the Y-axis is pile head displacement in millimeters (0 to 100) (0 to 3.94 inches). Davisson's lines are included for reference. Based on Davisson's criteria, the prestressed concrete test pile shows the highest capacity at about 3,100 kilonewtons. Capacities for the FRP composite pile and the plastic composite pile are approximately 25 to 30 percent lower than for the prestressed concrete pile. An axial load of 2,600 kilonewtons applied to the FRP pile results in a 40-millimeter pile head displacement. A similar axial load applied to the plastic composite pile results in a pile head displacement of about 95 millimeters. An axial load of about 3,100 kilonewtons applied to the prestressed concrete pile results in a 20-millimeter pile head displacement. One kilonewton equals 225 poundforce.

Figure 140. Graph. Distribution of residual loads. This figure presents the distribution of residual loads for the three types of test piles. The X-axis is load in kilonewtons (0 to 350) and the Y-axis is depth in meters (0 to 20) (0 to 66 feet). The chart compares the residual loads locked in the three test piles after driving. The chart indicates that the prestressed concrete pile retained the lowest load of about 120 kilonewtons near the 10-meter (32.8-foot) depth, and the plastic composite pile retained the highest load of about 320 kilonewtons near the 14-meter (46-foot) depth. The load retained by the FRP composite pile, about 140 kilonewtons near the 12-meter (39-foot) depth, was between the plastic pile and the prestressed concrete pile. One kilonewton equals 225 poundforce.

Figure 141. Graph. Distribution of residual stresses. This figure presents the distribution of residual stresses for the three types of test piles. The X-axis is residual stress in kilopascals (negative 40 to positive 40) and the Y-axis is depth in meters (0 to 20) (0 to 66 feet). The chart compares the residual stresses locked in the three test piles after driving. The residual stresses for the prestressed concrete pile and the FRP composite pile range narrowly between positive and negative 9 kilopascals. The residual stresses calculated for the plastic pile vary widely between about negative 40 and positive 30 kilopascals. One kilopascal equals 0.145 poundforce per square inch.

Figure 142. Graph. Load distribution for the three test piles at the Davisson failure loads. This figure presents load transfer diagrams for the three test piles to the ground. The X-axis is load in kilonewtons (0 to 3,500) and the Y-axis is depth in meters (0 to 20) (0 to 66 feet). For the plastic pile, the load transfer curve begins with a load of approximately 600 kilonewtons at a depth of approximately 18 meters (59.04 feet) and rises to a load of approximately 2,100 kilonewtons at 0 meters (0 feet). For the FRP pile, the load transfer curve begins with a load of approximately 600 kilonewtons at a depth of approximately 18 meters (59.04 feet) and rises to a load of approximately 2,250 kilonewtons at 0 meters (0 feet). For the prestressed concrete pile, the load transfer curve begins with a load of approximately 600 kilonewtons at a depth of approximately 18 meters (59.04 feet) and rises to a load of approximately 3,100 kilonewtons at 0 meters (0 feet). One kilonewton equals 225 poundforce.

Figure 143. Graph. Mobilized average unit shaft resistance-toe resistance relationships for the three test piles. This figure presents the relationship between the average unit shaft resistance and the mobilized toe resistance calculated for the three types of test piles. The X-axis is mobilized toe resistance in kilopascals (0 to 5,000) and the Y-axis is mobilized average unit shaft resistance in kilopascals (0 to 70). The chart indicates that the average unit shaft resistance decreases in order from the prestressed concrete pile, to the plastic composite pile, to the FRP composite pile, while the toe resistance increases in the same order from the prestressed concrete pile, to the plastic composite pile, to the FRP composite pile. One kilopascal equals 0.145 poundforce per square inch.

Figure 144. Graph. Apparent strength gain with time in the three test piles. This figure presents the apparent gain in strength with time after the end of initial driving for each of the three types of test piles. The X-axis is time after the end of initial driving in days (0 to 18) and the Y-axis is computed or measured capacity in kilonewtons (0 to 3,500). The prestressed concrete pile showed the greatest apparent gain in strength, from about 200 to 3,000 kilonewtons, over a period of 12 days. Over 11 days, the apparent gain for the FRP pile was from about 1,700 to 2,300 kilonewtons. The plastic pile increased from about 500 to 2100 kilonewtons over a period of 16 days. One kilonewton equals 225 poundforce.

Figure 145. Graphs. Deformed shapes of piles at different lateral loads. This figure consists of three graphs, one for each type of test pile, presenting the deformed shapes of the piles at six different static lateral loads and the corresponding displacement near the top of the pile. The X-axis on each chart is deflection in millimeters (0 to 120) (0 to 4.72 inches) and the Y-axis on each chart is depth below the top of the pile measured in meters (0 to 18) (0 to 59 feet). Ground level and the location of the bottom of the test pit are also indicated on the charts. All three charts reveal that increased lateral load also increases lateral deformation. Chart A for the prestressed concrete pile indicates that lateral deformation is insignificant from about 6 meters (19.7 feet) below the bottom of the test pit. Chart B for the FRP composite pile suggests that lateral deformation is insignificant from about 3 meters (10.8 feet) below the bottom of the test pit. Chart C for the plastic pile indicates that lateral deformation is unimportant from about 3.5 meters (11.5 feet) below the bottom of the test pit.

Figure 146. Graph. Measured lateral deflections at ground surface for the three test piles. This figure presents the relationship between the measured lateral load and the lateral deflection at the ground surface for each of the three test piles. The X-axis is lateral deflection at the ground surface measured in millimeters (0 to 120) (0 to 4.72 inches), and the Y-axis is measured lateral load in units of kilonewtons (0 to 350). Overall the chart indicates that the prestressed concrete pile and the FRP composite pile show similar load-deflection responses, and that the plastic composite pile exhibits the greatest deformation. At a lateral load of 200 kilonewtons, the lateral deflections measured were 33, 35, and 70 millimeters (1.37, 1.29, and 2.73 inches) for the FRP composite pile, the prestressed concrete pile, and the plastic composite pile, respectively. One kilonewton equals 225 poundforce.

Figure 147. Illustration. Load transfer in an axially loaded pile. This figure depicts the force components acting on a pile and that are used to estimate its ultimate axial load. The figure presents the equation: Q subscript T, which is the ultimate axial load capacity, equals Q subscript small S, which is the ultimate shaft capacity, plus Q subscript small B, which is the total ultimate tip load at the base of the pile, minus W subscript small P, which represents the weight of the pile. The forces Q subscript small S and Q subscript small B, needed to estimate the ultimate axial load capacity of a pile, along with the equations used to calculate each term, are shown acting on a pile, which is presented vertically in place below the ground surface.

Figure 148. Graphs. Interpreted average CPT and SPT design profiles for Route 351 test site. This figure consists of two side-by-side graphs presenting the interpreted average cone penetrometer test (CPT) and standard penetration test (SPT) profiles, which were used to predict the axial capacities of the three piles. The X-axis of the CPT profile is cone resistance in megapascals (0 to 14) and the X-axis of the SPT profile is the field N-value in blows per 300 millimeters (11.8 inches). The Y-axis for both charts is depth in meters (0 to 20) (0 to 66 feet). The interpreted average results indicate that cone resistance generally increases from about 8 to 14 megapascals in the 4- to 13-meter- (13- to 43-foot-) depth interval. Between 14 and 16 meters, a depth interval consisting of sandy clay soil, the cone resistance is minimized to about 2.5 to 5 megapascals. The interpreted average SPT profile shows stepwise increases in the field N-values with increased depth. Between the surface and 4 meters (13 feet), the average field N-value is 5. Between 4 and 11 meters (13 and 36 feet), the average N-value is 10. At a depth interval of 11 to 15 meters (36 to 49 feet), the average field N-value is 22. Between 15 and 20 meters (49 and 66 feet), the average N-value drops to 16. The conversion factor for megapascals is 1 megapascal equals 145 poundforce per square inch.

Figure 149. Graph. Accuracy of Nordlund's method predictions using small delta values from Nordlund's charts. For each of the three types of test piles,this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using Nordlund's charts to obtain the interface friction angles. The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the tip capacities are underpredicted for all test piles. The capacities for the FRP and plastic composite piles are underestimated by the largest percentage. The data also show that the shaft capacities are overpredicted for the three test piles. The shaft capacities for the FRP and plastic composite test piles are overpredicted by the largest percentage. The predicted total pile capacity for all test piles is slightly higher than the total measured capacity.

Figure 150. Graph. Accuracy of Nordlund's method predictions using small delta values from interface shear tests. For each of the three types of test piles,this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using the interface friction angles obtained from interface shear tests. The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the tip capacities are underpredicted for all test piles. The capacities for the FRP and plastic composite piles are underestimated by the largest percentage. The data also show that the shaft capacities are overpredicted for the three test piles. The shaft capacities for the FRP and plastic composite test piles are overpredicted by the largest percentage. The predicted total pile capacity for all test piles is higher than the total measured capacities.

Figure 151. Graph. Accuracy of API method predictions using lower-case delta values from table 36. For each of the three types of test piles,this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using the API-recommended soil-pile interface friction angles (table 36). The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the tip capacities using this API method are generally similar to the measured values, with the tip capacity for the prestressed concrete pile just slightly overpredicted, and the capacities for the plastic and FRP composite piles just slightly underpredicted. The data also show that the shaft capacities are lower than the measured values for all of the test piles. The shaft capacity for the prestressed concrete pile is underpredicted by the largest percentage. The total pile capacity for all test piles is underpredicted using the API-recommended interface friction angles.

Figure 152. Graph. Accuracy of API method predictions using lower-case delta values from interface shear tests. For each of the three types of test piles,this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using the interface friction angles determined from the shear interface tests. The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the tip capacities using this API method are generally similar to the measured values, with the tip capacity for the prestressed concrete pile just slightly overpredicted, and the capacities for the plastic and FRP composite piles just slightly underpredicted. For both the FRP and plastic composite piles, the predicted shaft capacities are very similar to the measured values. However, the predicted shaft capacity for the prestressed concrete pile is substantially lower than the measured value. Total pile capacity for the two composite test piles was nearly the same as the measured capacities.

Figure 153. Graph. Accuracy of LCPC method predictions using steel pile assumption. For each of the three types of test piles,this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using the LCPC method with the assumption that the test pile is steel. The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the predicted tip capacities using the LCPC "steel pile" assumption are higher than the measured capacities, with the tip capacity for the prestressed concrete pile substantially overpredicted by this method. The predicted shaft capacities are lower than the measured shaft capacities for all three piles. The shaft capacity calculated for the prestressed concrete pile is substantially lower than the measured capacity. Total predicted pile capacity for all test piles is less than the measured total capacities.

Figure 154. Graph. Accuracy of LCPC method predictions using concrete pile assumption. For each of the three types of test piles,this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using the LCPC method with the assumption that the test pile is concrete. The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the predicted tip capacities using the LCPC "concrete pile" assumption are higher than the measured capacities, with the tip capacity for the prestressed concrete pile substantially overpredicted by this method. The predicted shaft capacities for the two composite piles are higher than the measured capacities. For the concrete pile, the predicted shaft capacity is only slightly lower than the measured capacity. Total predicted pile capacity for all test piles is greater than the measured total capacities.

Figure 155. Graph. Accuracy of IC method predictions using lower-case delta values from interface shear tests. For each of the three types of test piles, this figure presents the ratio of the predicted pile capacity to the measured pile capacity for the pile tip, the pile shaft, and the total pile. The predicted capacities were calculated using the Imperial College method, which uses the average interpreted design CPT tip resistance profile (figure 148) together with the interface friction angle and surface roughness data. The X-axis is the pile capacity component (total, shaft, tip) and the Y-axis is the ratio of predicted capacity to measured capacity (0 to 2). Data on the chart indicate that the predicted tip capacities are generally in agreement with the measured capacities, with the tip capacity for the FRP composite pile slightly lower than measured, and the plastic composite pile and concrete prestressed concrete pile slightly higher than measured. The predicted shaft capacities for the two composite piles are slightly higher than the measured capacities. For the concrete pile, the predicted shaft capacity is lower than the measured capacity. Total predicted pile capacity for the prestressed concrete pile is also below the measured capacity; for the FRP pile and the plastic pile, the total predicted pile capacity is slightly above the measured capacity.

Figure 156. Illustration. Idealized model used in T-Z load transfer analyses. This figure depicts the elements of the load-transfer (T-Z) method used to study load-transfer behavior of a single axially loaded pile. It shows a pile in place below the ground surface, with the axial loads at the top and bottom of the pile and the side shear stress acting on the pile. Alongside this illustration of a pile is a set of vertically arranged discrete nonlinear springs, which are used in the load-transfer model to describe the resistance from the soil and end bearings. Alongside the column of springs are two small graphs describing the load-transfer relationships. The X-axis for the top graph is vertical displacement (Z), and the Y-axis is vertical shear (T). The graph is labeled T-Z curve for side springs. The X-axis for the bottom graph is vertical toe displacement (Z), and the Y-axis is vertical force (Q subscript small B). The graph is labeled Q subscript small B-Z curve for toe spring. Both curves show a logarithmic relationship.

Figure 157. Graph. Pile tip load-pile tip displacement curve (Q-Z) (API 1993). This figure presents the API load-transfer curve used to predict pile settlement. The X-axis is the ratio of the vertical toe displacement (Z) to the pile diameter (D) on a scale of 0 to 0.25. The Y-axis is the ratio of the mobilized end-bearing capacity (Q) to the ultimate end-bearing load (Q subscript small B) on a scale of 0 to 1.2. The graph, which shows a logarithmic relationship, also includes a small table presenting five values for each set of ratios for Q over Q subscript small B and for Z over D. The graph indicates that the ultimate end-bearing load is mobilized at a pile tip displacement equal to 10 percent of the pile diameter.

Figure 158. Graph. Maximum shear stress distribution along pile shaft, according to API (1993). This figure presents the distribution of the API-recommended maximum shear stresses along the pile shaft for each of the three test piles. The X-axis is shaft maximum shear stress in kilopascals (0 to 80) and the Y-axis is depth in meters (0 to 20) (0 to 66 feet). Between a depth of about 1 meter (3.3 feet) and 11 meters (36 feet), the maximum shear stresses for all three piles increase linearly from about 5 kilopascals to 67 kilopascals. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 159. Graph. Settlement predictions for the prestressed concrete pile using API (1993). This figure compares measured load-settlement responses for the prestressed concrete pile with predictions made using the API load-transfer curves, with and without a limit on shaft shear stress. The X-axis is applied axial load in kilonewtons (0 to 3,500) and the Y-axis is the pile head displacement in meters (0 to 0.03) (0 to 0.098 feet). The graph shows that the predicted responses agree well with the measured responses for the initial half of the settlement curve, to a load of about 2,250 kilonewtons. Above this value, both API prediction methods underpredict the measured pile capacity. The lower predicted capacity was derived from the API load-transfer curve using a limiting shaft shear stress. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 160. Graph. Settlement predictions for the FRP pile using API (1993). This figure compares measured load-settlement responses for the FRP composite pile with predictions made using the API load-transfer curves, with and without a limit on shaft shear stress. The X-axis is applied axial load in kilonewtons (0 to 3,000) and the Y-axis is the pile head displacement in meters (0 to 0.1) (0 to 0.33 feet). The graph shows that the predicted responses agree well with the measured responses for the initial half of the settlement curve, to a load of about 1,800 kilonewtons. However, both API methods underpredict the pile capacity measured in the field. The lower predicted capacity was derived from the API load-transfer curve limiting the shaft shear stress. The prediction based on the API load-transfer curve with no shaft shear stress cutoff better approximates the measurements made in the field. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 161. Graph. Settlement predictions for the plastic pile using API (1993). This figure compares measured load-settlement responses for the plastic composite pile with predictions made using the API load-transfer curves, with and without a limit on shaft shear stress. The X-axis is applied axial load in kilonewtons (0 to 3,000) and the Y-axis is the pile head displacement in meters (0 to 0.1) (0 to 0.33 feet). The chart show that the predicted responses agree fairly well with the measured responses for the initial half of the settlement curve, to a load of about 1,700 kilonewtons. The curve representing the API method that limits shaft shear stress indicates that, at about 1,700 kilonewtons, large settlements are predicted sooner than the field observations. The API prediction method that does not limit the shaft shear stresses provides data that agree well with the field observations over the entire test period. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 162. Graph. Normalized T-Z curves according to API (1993) and Vijayvergiya (1977). This figure compares the normalized Vijayvergiyaand the APIload-transfer curves. The X-axis is local shaft displacement in meters (0 to 0.018) (0 to 0.059 feet) and the Y-axis is the normalized shaft shear stress from 0 to 1.2. The primary difference between the two curves is the shape. The normalized Vijayvergiya load-transfer curve is logarithmic and the API load transfer is a two-part response, initially linear, then flat across the range of shaft displacement at a normalized shaft shear stress of 1.

Figure 163. Graph. Normalized Q-Z curves according to Vijayvergiya (1977) and API (1993). This figure compares the normalized Vijayvergiya and the API pile tip load versus pile tip displacement (Q-Z) curves. The X-axis is pile tip displacement in meters (0 to 0.1) (0 to 0.33 feet) and the Y-axis is the normalized pile tip stress from 0 to 1.2. The Vijayvergiya Q-Z curve appears to mobilize the full end-bearing capacity at a smaller displacement than the API curve for a 610-millimeter- (24-inch-) diameter pile.

Figure 164. Graph. Maximum shear stress distributions used in predictions using Vijayvergiya (1977). This figure presents the distribution of Vijayvergiya-recommended maximum shear stresses along the pile shaft for each of the three test piles using two different coefficients of lateral earth pressure (K). The X-axis is shaft maximum shear stress in kilopascals (0 to 150) and the Y-axis is depth in meters (0 to 18) (0 to 59 feet). For both coefficients of lateral earth pressure shown, shear stress along the pile shaft increases with depth for the three test piles. At a coefficient of lateral earth pressure equal to 1, the shaft shear stress increases linearly from about 12 kilopascals to about 105 kilopascals over the 18-meter (59-foot) pile depth. At a coefficient of lateral earth pressure equal to 1.25, which might be encountered in medium dense to dense sand, the shear stress for all pile types is 20 to 30 percent higher. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 165. Graph. Settlement predictions for the concrete pile using Vijayvergiya (1977). This figure compares measured load-settlement responses for the prestressed concrete pile with predictions made using the Vijayvergiya load-transfer curves, considering two coefficients of lateral earth pressure, K equals 1 and K equals 1.25. The X-axis is applied axial load in kilonewtons (0 to 3,500) and the Y-axis is the pile head displacement in meters (0 to 0.03) (0 to 0.098 feet). The chart suggests that the predicted responses agree well with the measured responses for the initial part of the settlement curves for each coefficient of lateral earth pressure, to a load of about 1,000 kilonewtons. For loads exceeding 1,000 kilonewtons, the Vijayvergiya-based predictions are generally higher than the measured responses. An exception is when K equals 1 and the load exceeds approximately 2,700 kilonewtons; then the predicted pile head displacement falls below the measured displacement. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 166. Graph. Settlement predictions for the FRP pile using Vijayvergiya (1977). This figure compares measured load-settlement responses for the FRP composite pile with predictions made using the Vijayvergiya load-transfer curves and a coefficient of lateral earth pressure of K equals 1. The X-axis is applied axial load in kilonewtons (0 to 3,000) and the Y-axis is the pile head displacement in meters (0 to 0.04) (0 to 0.13 feet). The chart suggests that the load-settlement predictions, based on the Vijayvergiya method and for a coefficient of lateral earth pressure equal to 1, agree fairly well with the field measurements on the FRP pile. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 167. Graph. Settlement predictions for the plastic pile using Vijayvergiya (1977). This figure compares measured load-settlement responses for the plastic composite pile with predictions made using the Vijayvergiya load-transfer curves and a coefficient of lateral earth pressure of K equals 1. The X-axis is applied axial load in kilonewtons (0 to 3,000) and the Y-axis is the pile head displacement in meters (0 to 0.1) (0 to 0.33 feet). The chart suggests that the load-settlement predictions, based on the Vijayvergiya method and for a coefficient of lateral earth pressure equal to 1, are higher than the field measurements on the plastic pile. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 168. Illustrations. Concentric cylinder model for settlement analysis of axially loaded piles (adapted from Randolph and Wroth 1978). This figure consists of two illustrations that graphically present the terms of a theoretical description of soil deformation around a pile shaft. This description forms the basis for the concentric cylinder model to describe axially loaded piles. The first illustration (A) is a loaded tubular pile surrounded by two concentric cylinders, and showing the load and stress forces as arrows. The concentric cylinders in shear, which theoretically describe the soil deformation process, are shown enclosing a ring of soil, a wedge of which is enlarged and expanded in the second illustration (B). Illustration (B) presents the stress and displacement field for the soil wedge and shows the stresses due to vertical and lateral loading on the wedge as arrows in the direction of the stresses. Many of the stresses are described in compound terms.

Figure 169. Graph. Linear T-Z Curve obtained using Randolph and Wroth (1978). This figure presents the T-Z load-transfer curve obtained from the concentric cylinder model for predicting pile settlement. The curve assumes a linear and elastic soil. The X-axis is local shaft displacement in meters (0 to 0.0025) (0 to 0.0082 feet) and the Y-axis is the normalized shaft shear stress (0 to 1.2). The curve is linear from 0 to 1 on the normalized shaft shear scale and between a shaft displacement of 0 and 0.001 meters (0 and 0.0033 feet). The slope of the line is described by equation 41, which is K subscript S equals G divided by the product of R subscript O times the natural logarithm of the quotient of R subscript M divided by R subscript O.

Figure 170. Graph. Linear Q subscript B-Z curve obtained using Boussinesq's theory. This figure presents the Q-Z load-transfer curve for the pile tip based on Boussinesq's equation, assuming a linear and elastic soil. The X-axis is local shaft displacement in meters (0 to 0.02) (0 to 0.066 feet) and the Y-axis is the normalized shaft shear stress (0 to 1.2). The curve is linear for a normalized pile tip load between 0 and 1, and a pile tip displacement of 0 to 0.006 meters (0 and 0.02 feet). The slope of the line corresponds to the stiffness coefficient of the pile tip spring, which is described by equation 44, which is K subscript base equals the quotient of the product of 4 times G times R subscript O, divided by the sum of 1 minus nu.

Figure 171. Graph. Hyperbolic T-Z curve. This figure compares the nonlinear hyperbolic model with the linear elastic curve for predicting pile settlement. The X-axis is the local shaft displacement in meters (0 to 0.004) (0 to 0.013 feet) and the Y-axis is normalized shaft shear stress (0 to 1.2). The figure clearly demonstrates the difference in shape between the two curves. The linear elastic T-Z curve is entirely linear and the hyperbolic T-Z curve is basically logarithmic in shape, with only the initial slope corresponding to the linear response. The slope of the linear portion of the hyperbolic T-Z curve is described by an equation almost identical to equation 41, which is K subscript S equals G divided by the product of R subscript O times the natural logarithm of the quotient of R subscript M divided by R subscript O.

Figure 172. Graph. Variation of secant shear modulus for different hyperbolic-type models. This figure compares the relationship between the normalized secant shear modulus and the normalized shear stress for three hyperbolic models and for a modified hyperbolic model. The X-axis is normalized shear stress (0 to 1) and the Y-axis is normalized secant shear modulus (0 to 1). This relationship for the three hyperbolic models, with failure ratios or slopes of 1, 0.6, and 0, produce straight lines. In contrast, the relationship between the normalized secant shear modulus and the normalized shear stress for the Fahey and Carter (1993) modified hyperbolic model, which alters the empirical curve-fitting parameters, results in an exponentially decreasing curve.

Figure 173. Graph. Theoretically derived T-Z curve using concentric cylinders, and the modified hyperbola from Fahey and Carter (1993). This figure compares three theoretically derived T-Z load-transfer curves. The first two-the linear elastic response and the hyperbolic with R subscript F equal to 1.0-were depicted in figure 171. The third is the Fahey and Carter (1993) hyperbola and is based on equation 55 in the text. The X-axis is the local shaft displacement in meters (0 to 0.006) (0 to 0.02 feet) and the Y-axis is the normalized shaft shear stress (0 to 1.2). The Fahey and Carter hyperbola is a logarithmic curve, but one that deviates from a straight line faster than the hyperbolic curve from figure 171.

Figure 174. Graph. Theoretically derived Q-Z curve using Boussinesq's theory and the modified hyperbola from Fahey and Carter (1993). This figure compares three theoretically derived Q-Z load-transfer curves. The linear elastic response, derived from the Boussinesq model to describe load-transfer at the pile tip, is compared with the nonlinear hyperbolic model and the modified hyperbolic model. The X-axis is the normalized pile tip displacement (0 to 0.1) and the Y-axis is the normalized pile tip load (0 to 1.2). For this comparative relationship, the hyperbolic models result in logarithmic curves, and the Boussinesq's linear elastic model results in a straight line. The three curves initially overlap, and the slopes of the lines in the overlap portion are given by an equation similar to equation 44, which is K superscript base subscript O equals the quotient of the product of 4 times G subscript O times R subscript O, divided by the sum of 1 minus nu.

Figure 175. Graph. Route 351 initial shear modulus profile from CPT correlations. This figure compares depth profiles of initial shear modulus estimated by three different methods and using the design CPT profile obtained from the bridge project site (figure 148). The X-axis is initial shear modulus in megapascals (0 to 150) and the Y-axis is depth in meters (0 to 20) (0 to 66 feet). The three profiles, determined according to equations of Rix and Stokoe (1991), Baldi, et al. (1989), and Chow (1996), are similar across the depth profile and vary slightly in magnitude. The method of Baldi, et al.(1989), results in the highest values for initial shear modulus. The lowest values are derived from the Chow (1996) equation. The general trend is an increase in initial shear modulus with depth. For all three profiles, peaks are shown at about 13 and 18 meters (43 and 59 feet). The conversion factor for megapascals is 1 megapascal equals 145 poundforce per square inch.

Figure 176. Graph. Settlement predictions for the concrete pile using theoretically derived transfer curves. For the prestressed concrete test pile,this figure compares predicted load-settlement curves with the load-settlement curve derived from field measurements. The predictions are based on the proposed theoretically derived load-transfer curves and are calculated for two coefficients of horizontal earth pressure. The X-axis is applied axial load in kilonewtons (0 to 3,500) and the Y-axis is the pile head displacement in meters (0 to 0.03) (0 to 0.098 feet). The figure indicates that the predicted load-settlement curve, using 1.25 as the coefficient of horizontal earth pressure, agrees fairly well with the curve derived from field measurements. The analysis conducted with 1.0 as the coefficient of horizontal earth pressure predicts the initial slope of the load-settlement curve and ultimate pile capacity well, but shows a slightly stiffer response for loads between 1,500 and 2,900 kilonewtons. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 177. Graph. Settlement prediction for the FRP pile using theoretically derived transfer curves. For the FRP composite test pile, this figure compares a predicted load-settlement curve with the load-settlement curve derived from field measurements. The prediction is based on the proposed theoretically derived load-transfer curves and is calculated using 1.0 as the coefficients of horizontal earth pressure. The X-axis is applied axial load in kilonewtons (0 to 3,000) and the Y-axis is the pile head displacement in meters (0 to 0.04) (0 to 0.13 feet). The figure indicates that the predicted load-settlement curve agrees very well with the curve derived from field measurements. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 178. Graph. Settlement predictions for the plastic pile using theoretically derived transfer curves. For the plastic composite test pile,this figure compares predicted load-settlement curves with the load-settlement curve derived from field measurements. The predictions are based on the proposed theoretically derived load-transfer curves and are calculated for two coefficients of horizontal earth pressure. The X-axis is applied axial load in kilonewtons (0 to 3,000) and the Y-axis is the pile head displacement in meters (0 to 0.1) (0 to 0.33 feet). The figure indicates that the predicted load-settlement curve, using 0.9 as the coefficient of horizontal earth pressure, approximates the field-measured response fairly well until the load reaches about 2,250 kilonewtons. The analysis using 1 as the coefficient of lateral earth pressure overpredicts the measured response. Neither analysis approximates the slope of the second half of the curve based on measured values. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure. 179. Illustration. Laterally loaded pile problem. This figure consists of an illustration of a loaded pile below the ground surface (A), correlated with four typical depth profiles for (B) the net soil reaction, (C) the pile deflection, (D) the slope, and (E) the bending moment. To illustrate the derivation of the differential equation for laterally loaded piles, the loaded pile (A) is divided into infinitesimal elements, DX, along the distance of the pile. One of these elements is enlarged to show the forces acting on it. Representative profiles for the net soil reaction (B), pile deflection (C), and slope (D) are shown decreasing with depth. The bending moment (E) is shown increasing in value from the top of the pile, reaching a maximum value just below the surface, and then decreasing with depth.

Figure 180. Illustration. Distribution of stresses against a pile before and after lateral loading (adapted from Reese and Van Impe 2001). This figure consists of two sets of drawings, each set showing a longitudinal view and a cross section of a pile. The two sets of drawings depict the distribution of stresses acting on a cylindrical pile before lateral loading (A) and after lateral loading (B). Before a lateral load is applied to the pile (A), the longitudinal view shows that the pile is vertically aligned, and the cross section reveals that the stresses along a horizontal plane are uniform and normal to the pile. After lateral loading, the longitudinal view shows that the pile is deflected, and its cross section reveals that the soil stresses along a horizontal plane are not uniform, but have both normal and tangential distributions.

Figure 181. Graphs. Typical P-Y curve and resulting P-Y modulus (Reese and Van Impe 2001). This figure consists of two charts showing the typical relationship between pile deflection and soil reaction (A), and between pile deflection and the P-Y modulus. The X-axis for both charts is a unit of length. The Y-axis for chart A is in units of force divided by length. For chart B, the Y-axis is in units of force divided by the length squared. Chart A of the figure illustrates the typical logarithmic P-Y curve showing the initial slope, the ultimate soil resistance value, and the P-Y modulus or the reaction modulus. The equation for the soil reaction is also shown, relating the pile deflection to the P-Y modulus. Chart B presents the graphic relationship between the pile deflection and the reaction modulus. The shape of the curve for this relationship is inversely exponential.

Figure 182. Illustration. Schematic showing the influence of shape of cross section of pile on the soil reaction P (adapted from Reese and Van Impe 2001). This figure illustrates the influence of three pile shapes on soil resistance. Cross sections for a cylindrical pile, a square-column pile, and a rectangular column with a small width versus length, are shown with stresses affecting the resistance. The soil resistance is greatest for the square-column pile, less for the cylindrical pile, and least for the board-shaped rectangular-column pile.

Figure 183. Graph. Elements of a characteristic P-Y curve for sand based on recommendations by Reese, et al. (1974). This figure presents a typical P-Y curve for sand and identifies the elements defining the curve. The X-axis is the deflection in units of length and the Y-axis is the soil reaction in units of force divided by length. Elements identified on the curve include the initial P-Y modulus, which defines the initial linear part of the curve to point A; point B, the coordinates of which are the pile width B divided by 60; point C, the coordinates of which are three times the pile width divided by 80, and the ultimate soil resistance, which defines the curve at point C and beyond. The section of the curve between points A and C is the transition zone to connect the ultimate soil resistance with the initial P-Y modulus line. This transitional section is divided into a parabolic section between points A and B, and a linear section between points B and C.

Figure 184. Illustration. Schematic showing P-Y model used for analysis of laterally loaded piles. This figure depicts the elements of the P-Y model for analysis of the lateral load transfer behavior of a single axially loaded pile. It shows a pile in place below the ground surface and the lateral load at the top of the pile. Alongside the pile is a set of vertically arranged discrete nonlinear springs, which are used in this model to describe the resistance from the soil. A small chart presenting the logarithmic relationship between pile deflection Y and soil-pile reaction P is also shown for two depths. The initial slope of the P-Y curve increases with increasing depth.

Figure 185. Graphs. In situ test data for the upper soils in the northern end of the test pile site. This figure consists of three adjacent graphs presenting the standard penetration test (SPT) profile, and the cone penetrometer test profiles for the tip resistance and sleeve friction measured in the upper 10 meters (33 feet) of soil at the northern part of the test pile site. The X-axis of the SPT profile is the field N-value (0 to 20). The X-axis of the CPT tip resistance profile is cone resistance in bars (0 to 200) and the X-axis of the CPT sleeve friction profile is sleeve friction in bars (0 to 2). The Y-axis for all three charts is depth in meters (0 to 10) (0 to 33 feet). The SPT profile shows that the upper 1 meter (3.3 feet) of soil is silty fine sand (SM) fill, below which is loose to medium dense (SM) silty file sand. The field N-value is highest in the upper SM fill layer and then is relatively constant throughout the lower SM layer. The greatest cone tip resistance is shown near the surface and again appears relatively constant with depth throughout the upper 10 meters (33 feet). The sleeve friction is more variable, showing high values near the surface and then again between 5 and 10 meters (16.4 and 33 feet).

Figure 186. Graphs. In situ test data for the upper soils at the southern end of the test pile site. This figure consists of four adjacent graphs presenting the standard penetration test (SPT) profile, the cone penetrometer test profiles for the tip resistance and sleeve friction, and the flat dilatometer test (DMT) profile measured in the upper 10 meters (33 feet) of soil at the southern part of the test pile site. The X-axis of the SPT profile is the field N-value (0 to 20). The X-axis of the CPT tip resistance profile is cone resistance in bars (0 to 200) and the X-axis of the CPT sleeve friction profile is sleeve friction in bars (0 to 2). The X-axis of the DMT profile is the modulus in bars (0 to 100). The Y-axis for all charts is depth in meters (0 to 10) (0 to 33 feet). The SPT profile shows that the upper 0.5 meters (1.64 feet) of soil is silty sand (SM) fill, below which is a 1-meter- (3.3-foot-) thick layer of medium stiff sand, silty clay (CL), followed by medium dense silty sand (SM). The field N-value is around 7 in the upper layer, decreases to about 5 in the sandy clay, and increases gradually in the medium dense sand layer to a maximum of about 14. The greatest cone tip resistance, about 100 bars, is shown near the surface, then decreases to below 25 bars in the area of the clay layer, and increases to about 60 bars, where it is relatively constant with depth. The sleeve friction values are high near the surface, low in the clay layer, and high again in the medium dense sand. The profile for the DMT modulus appears more irregular, showing a peak at the interface between the surface layer and the sandy clay at about 0.5 meters (1.64 feet), a decrease in the clay and upper part of the medium dense sand (1.5 to 3 meters) (4.9 to 9.8 feet), and another spike at a depth of 3.5 meters (11.5 feet).

Figure 187. Graph. Initial P-Y modulus profile used to define default P-Y curves for LPILE analyses on the prestressed concrete pile. This figure presents the initial P-Y modulus profile used to generate the P-Y curves needed for input to the LPILE model. The X-axis is the initial P-Y modulus in units of meganewtons per meters squared (0 to 20) and the Y-axis is depth below the ground surface in meters (0 to 10) (0 to 33 feet). The relationship between the initial pile modulus and depth is linearly increasing. The slope of the line is also described in mathematical terms. The conversion factor for meganewtons is 1 meganewton equals 225,000 poundforce.

Figure 188. Graphs. Predicted versus measured lateral displacement profile for prestressed concrete pile (low lateral loads). This figure consists of two graphs that compare the predicted lateral pile deflection profiles for the prestressed concrete pile in sand. The pile deflections result from lateral loads of 51.2 and 97.3 kilonewtons (11,520 and 21,893 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 80) (0 to 3.15 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads, the graphs indicate that the deflected shapes measured for the prestressed concrete piles agree fairly well with the model predictions. The predicted deflection values, ranging from 0 to about 10 millimeters (0 to 0.39 inches) for the 51.2-kilonewton (11,520-poundforce) load, and from 0 to about 20 millimeters (0 to 0.79 inches) for the 97.3-kilonewton (21,893-poundforce) load are, however, slightly higher than the field measurements.

Figure 189. Graphs. Predicted versus measured lateral displacement profile for prestressed concrete pile (medium lateral loads). This figure consists of two graphs that compare the predicted lateral pile deflection profiles for the prestressed concrete pile in sand. The pile deflections result from lateral loads of 141.2 and 186.4 kilonewtons (31,770 and 41,940 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 80) (0 to 3.15 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads, the graphs indicate that the deflected shapes measured for the prestressed concrete piles agree very well with the model predictions. The lateral deflections range from 0 to 30 millimeters (0 to 1.18 inches) for the 141.2-kilonewton (31,770-poundforce) load and from 0 to 38 millimeters (0 to 1.5 inches) for the 186.4-kilonewton (41,940-poundforce) load.

Figure 190. Graphs. Predicted versus measured lateral displacement profile for prestressed concrete pile (high lateral loads). This figure consists of two graphs that compare the predicted lateral pile deflection profiles for the prestressed concrete pile in sand. The pile deflections result from lateral loads of 228 and 247.7 kilonewtons (51,300 and 55,733 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 80) (0 to 3.15 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads, the graphs indicate that the deflected shapes measured for the prestressed concrete piles agree reasonably well with the model predictions. The predicted deflection values, ranging from 0 to about 50 millimeters (0 to 1.97 inches) for the 228-kilonewton (51,300-poundforce) load, and from 0 to about 60 millimeters (2.36 inches) for the 247.7-kilonewton (55,733-poundforce) load are, however, slightly lower than the field measurements.

Figure 191. Graph. Calculated load-deflection curve for the prestressed concrete pile. This figure presents a comparison of the relationship between the lateral pile deflection at the ground surface and the applied lateral load, as computed by the LPILE soil model, with measurements made on the prestressed concrete pile in the field. The X-axis is lateral deflection at the ground surface in millimeters (0 to 60) (0 to 2.36 inches) and the Y-axis is lateral load in kilonewtons (0 to 300). The predicted values are shown for both a constant and a variable flexural stiffness. The graph suggests that, overall, the model-computed lateral pile deflection agrees fairly well with the field measurements made on the prestressed concrete pile. At lateral loads greater than approximately 180 kilonewtons, the model calculation using a constant flexural stiffness is slightly lower than the model prediction using a variable flexural stiffness. Below a lateral load of about 180 kilonewtons, similar results are predicted using either the variable or the constant flexural stiffness. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 192. Graph. Calculated load-slope curve for the prestressed concrete pile. This figure presents a comparison of the relationship between the pile head rotation at the ground surface and the applied lateral load, as computed by the LPILE soil model, with measurements made in the field. The X-axis is the pile head slope in radians (0 to 0.012) and the Y-axis is lateral load in kilonewtons (0 to 300). The predicted values are shown for both a constant and a variable flexural stiffness. The graph suggests that the soil model predicts slightly higher values for the pile head rotation but, overall, agrees fairly well with the field measurements made on the prestressed concrete pile. At lateral loads greater than approximately 170 kilonewtons, the model prediction using a constant flexural stiffness is slightly lower than the model prediction using a variable flexural stiffness. Below a lateral load of about 170 kilonewtons, similar results are predicted using either the variable or the constant flexural stiffness. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 193. Graph. Initial P-Y modulus profile used to define default P-Y curves for LPILE analyses on the FRP pile. This figure presents the initial P-Y modulus profile used to generate the P-Y curves needed for input to the LPILE model. The X-axis is the initial P-Y modulus in units of meganewtons per meters squared (0 to 60) and the Y-axis is depth below the ground surface in meters (0 to 10) (0 to 33 feet). The relationship between the initial pile modulus and depth is linearly increasing. The slope of the line is also described in mathematical terms. The conversion factor for meganewtons is 1 meganewton equals 225,000 poundforce.

Figure 194. Graphs. Predicted versus measured lateral displacement profile for FRP pile (low lateral loads). This figure consists of two graphs that compare the predicted lateral pile deflection profiles for the FRP composite pile in sand. The pile deflections result from lateral loads of 51.6 and 96 kilonewtons (11,610 and 21,600 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 80) (0 to 3.15 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads, the graphs indicate that the deflected shapes measured for the FRP composite pile agree well with the model predictions. The predicted deflection values, ranging from 0 to about 10 millimeters (0 to 0.39 inches) for the 51.6-kilonewton (11,610-poundforce) load, are slightly higher than the field measurements. Predicted deflection values for the 96-kilonewton (21,600-poundforce) load range from 0 to about 15 millimeters (0 to 0.59 inches) and agree very well with the field measurements.

Figure 195. Graphs. Predicted versus measured lateral displacement profile for FRP pile (medium lateral loads). This figure consists of two graphs that compare the predicted lateral pile deflection profiles for the FRP composite pile in sand. The pile deflections result from lateral loads of 144.8 and 186.1 kilonewtons (32,580 and 41,873 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 80) (0 to 3.15 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads the graphs indicate that the deflected shapes measured for the FRP composite pile agree very well with the model predictions. The predicted deflection values ranging from 0 to about 30 millimeters (0 to 1.18 inches) for the 144.8-kilonewton (32,580-poundforce) load, and from 0 to about 40 millimeters (0 to 1.57 inches) for the 186.1-kilonewton (41,873-poundforce) load.

Figure 196. Graphs. Predicted versus measured lateral displacement profile for FRP pile (high lateral loads). This figure consists of two graphs that compare the predicted lateral pile deflection profiles for the FRP composite pile in sand. The pile deflections result from lateral loads of 230.1 and 270.5 kilonewtons (51,773 and 60,863 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 80) (0 to 3.15 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads the graphs indicate that the deflected shapes measured for the FRP composite pile agree well with the model predictions. The predicted deflection values range from 0 to about 55 millimeters (0 to 2.16 inches) for the 230.1-kilonewton (51,773-poundforce) load and from 0 to about 62 millimeters (0 to 2.44 inches) for the 270.5-kilonewton (60,863-poundforce) load.

Figure 197. Graph. Calculated load-deflection curve for the FRP pile. This figure presents a comparison of the relationship between the lateral pile deflection at the ground surface and the applied lateral load, as computed by the LPILE soil model, with measurements made on the FRP composite pile in the field. The X-axis is lateral deflection at the ground surface in millimeters (0 to 60) (0 to 2.36 inches) and the Y-axis is lateral load in kilonewtons (0 to 300). The predicted values are shown for both a constant and a variable flexural stiffness. The graph suggests that the lateral pile deflection calculated by the soil model using a variable flexural stiffness agrees well with the field measurements made on the FRP composite pile. At lateral loads greater than 100 kilonewtons, the model prediction using a constant flexural stiffness is considerably lower than the model prediction using a variable flexural stiffness. Below a lateral load of about 100 kilonewtons, similar results are predicted using either the variable or the constant flexural stiffness. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 198. Graph. Calculated load-slope curve for the FRP pile. This figure presents a comparison of the relationship between the pile head rotation at the ground surface and the applied lateral load, as computed by the LPILE soil model, with measurements made on the FRP composite in the field. The X-axis is the pile head slope in radians (0 to 0.025) and the Y-axis is lateral load in kilonewtons (0 to 300). The predicted values are shown for both a constant and a variable flexural stiffness. The chart suggests that the lateral pile deflection calculated by the soil model using a variable flexural stiffness agrees fairly well with the field measurements made on the FRP composite pile. At lateral loads greater than 100 kilonewtons, the model prediction using a constant flexural stiffness is considerably lower than the model prediction using a variable flexural stiffness. Below a lateral load of about 100 kilonewtons, similar results are predicted using either the variable or the constant flexural stiffness. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 199. Graph. Initial P-Y modulus profile used to define default P-Y curves for LPILE analyses on the plastic pile. This figure presents the initial P-Y modulus profile used to generate the P-Y curves needed for input to the LPILE model. The X-axis is the initial P-Y modulus in units of meganewtons per meters squared (0 to 25) and the Y-axis is depth below the ground surface in meters (0 to 10) (0 to 33 feet). The relationship between the initial pile modulus and depth is linearly increasing. The slope of the line is also described in mathematical terms. The conversion factor for meganewtons is 1 meganewton equals 225,000 poundforce.

Figure 200. Graphs. Predicted versus measured lateral displacement profile for plastic pile (low lateral loads). This figure consists of two charts that compare the predicted lateral pile deflection profiles for the plastic composite pile in sand. The pile deflections result from lateral loads of 48.1 and 102.4 kilonewtons (10,823 and 23,040 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 140) (0 to 5.5 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads the graphs indicate that the deflected shapes measured for the plastic composite pile agree well with the model predictions. The predicted deflection values range from 0 to 20 millimeters (0 to 0.79 inches) for the 48.1-kilonewton (10,823-poundforce) load and from 0 to 40 millimeters (0 to 1.57 inches) for the 102.4-kilonewton (23,040 poundforce) load.

Figure 201. Graphs. Predicted versus measured lateral displacement profile for plastic pile (medium lateral loads). This figure consists of two charts that compare the predicted lateral pile deflection profiles for the plastic composite pile in sand. The pile deflections result from lateral loads of 140.1 and 183.1 kilonewtons (31,523 and 41,198 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 140) (0 to 5.5 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads the graphs indicate that the deflected shapes measured for the plastic composite pile agree well with the model predictions. The predicted deflection values range from 0 to 60 millimeters (0 to 2.36 inches) for the 140.1-kilonewton (31,523-poundforce) load and from 0 to 80 millimeters (0 to 3.15 inches) for the 183.1-kilonewton (41,198-poundforce) load.

Figure 202. Graphs. Predicted versus measured lateral displacement profile for plastic pile (high lateral loads). This figure consists of two charts that compare the predicted lateral pile deflection profiles for the plastic composite pile in sand. The pile deflections result from lateral loads of 230.8 and 275.3 kilonewtons (51,930 and 61,943 poundforce), with the lateral pile deflection profiles determined by field measurements. The X-axis on both graphs is the lateral pile deflection in millimeters (0 to 140) (0 to 5.5 inches). The Y-axis is depth below the top of the pile in meters (0 to 18) (0 to 59 feet). For the lateral loads the graphs indicate that the deflected shapes measured for the plastic composite pile agree well with the model predictions. The predicted deflection values range from 0 to 100 millimeters (0 to 3.94 inches) for the 230.8-kilonewton (51,930-poundforce) load and from 0 to 120 millimeters (0 to 4.72 inches) for the 275.3-kilonewton (61,943-poundforce) load.

Figure 203. Graph. Calculated load-deflection curve for the plastic pile. This figure presents a comparison of the relationship between the lateral pile deflection at the ground surface and the applied lateral load, as computed by the LPILE soil model, with measurements made on the plastic composite pile in the field. The X-axis is lateral deflection at the ground surface in millimeters (0 to 100) (0 to 3.94 inches) and the Y-axis is lateral load in kilonewtons (0 to 300). For this pile, the constant and variable flexural stiffness are the same. The graph suggests that the lateral pile deflection calculated by the soil model agrees well with the field measurements made on the plastic composite pile. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 204. Graph. Calculated load-slope curve for the plastic pile. This figure presents a comparison of the relationship between the pile head rotation at the ground surface and the applied lateral load, as computed by the LPILE soil model, with measurements made on the plastic composite pile in the field. The X-axis is the pile head slope in radians (0 to 0.035) and the Y-axis is lateral load in kilonewtons (0 to 300). The relationship between the pile head slope and the lateral load is linear. For this pile, the constant and variable flexural stiffness is the same. The graph suggests that the lateral pile deflection calculated by the soil model agrees well with the field measurements made on the plastic composite pile. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 205. Illustration. Location of instrumented production piles at the Route 351 Bridge. This figure consists of two line drawings of the new Route 351 bridge. The first illustration is an elevation view of the proposed bridge, showing the superstructure supported by 12 piers and 2 end abutments. This drawing shows that the first four spans on the west side of the bridge are steel plate girder spans, and the remaining spans are prestressed concrete beam spans. The second drawing shows the 12 piers, the 2 end abutments, their orientation, and the distance between each pier. Each pier, except number 2, is shown supported by eight prestressed concrete piles. The orientation of each of the 1-to-4 battered piles is also indicated by an arrow on the pier schematics. From west to east, pier number 1 is located 15.2 meters (49.9 feet) from the west end abutment. Pier number 2 is located 26.6 meters (87.3 feet) from pier number 1. Pier number 3 is located 25.4 meters (83.3 feet) from pier number 2. Pier number 4 is located 21.29 meters (69.8 feet) from pier number 3. Piers 5 through 12 are each located 15.49 meters (50.8 feet) apart and pier number 12 is also 15.49 meters (50.8 feet) from the east end abutment. The drawing also shows the location of an instrumented prestressed concrete pile in the center of pier number 10 and an instrumented FRP composite pile in the center of pier number 11.

Figure 206. Illustration. Load transfer instrumentation layout for prestressed concrete production pile. This figure consists of two drawings depicting the vertical layout and a cross section of the instrument-fitted prestressed concrete production pile. The vertical diagram shows the locations of 18 vibrating wire sister-bar strain gages along the 22-meter- (72-foot-) long pile. A pair of sister bars, on opposite sides of the pile, is located at seven intervals along the pile. Two additional gages are shown at the top and also at the bottom of the test pile. Four sister bars (SB-1) are located at a depth of 2.31 meters (7.6 feet) from the top of the pile. Two sister bars (SB-2) are shown 3.69 meters (12 feet) below SB-1. Three meters (9.8 feet) below SB-2 are two additional sister bars (SB-3). A pair of sister bars (SB-4) is located 2.99 meters (9.8 feet) below SB-3. Another pair of sister-bar gages (SB-5) is located 3.01 meters (9.9 feet) below SB-4. Three meters (9.8 feet) below SB-5 is another pair of sister bars (SB-6). The last set of four sister bars (SB-7) is located 2.68 meters (8.8 feet) below SB-6 and 1.32 meters (4.3 feet) above the bottom of the production pile. The cross section diagram of the instrumented production pile shows a pair of sister-bar gages, each mounted 95 millimeters (3.74 1inches) in from the outer wall and on opposite sides of the 610-millimeter- (24-inch-) square column pile.

Figure 207. Illustration. Instrumentation layout for FRP composite production pile. This figure consists of two drawings depicting the vertical layout and one showing a cross section of the instrument-fitted FRP composite production pile. The first vertical diagram shows the locations of 18 vibrating wire sister-bar strain gages along the 21.2-meter- (69.5-foot-) long pile. A pair of sister bars, on opposite sides of the pile, is located at seven intervals along the pile. Two additional gages are shown at the top and also at the bottom of the test pile. Four sister bars (SB-1) are located at a depth of 2.3 meters (7.55 feet) from the top of the pile. Two sister bars (SB-2) are shown 3.7 meters (12.1 feet) below SB-1. Three meters (9.8 feet) below SB-2 are two additional sister bars (SB-3). A pair of sister bars (SB-4) is located 3.0 meters (9.8 feet) below SB-3. Another pair of sister-bar gages (SB-5) is located 3.0 meters (9.8 feet) below SB-4. Three meters (9.8 feet) below SB-5 is another pair of sister bars (SB-6). The last set of four sister bars (SB-7) is located 2.7 meters (8.86 feet) below SB-6. The second vertical diagram of the pile illustrates the locations of instruments bonded to the inside FRP shell. This diagram reveals the locations of 12 foil strain gages, mounted on opposite sides of the pile and bonded to the internal wall of the FRP tube at each of 6 levels. Of these 12 gages, 8 are shown mounted in the axial direction and 4 are shown mounted in the hoop direction. The first set of axially mounted strain gages (IG-1) is located 1.14 meters (3.74 feet) from the top of the pile. The next pair (IG-2) of axially mounted gages is located 0.44 meters (1.44 feet) below IG-1. The third set of strain gages (IG-3) is mounted in the hoop direction at a depth of 0.33 meters (1.08 feet) below IG-2. The next pair (IG-4) is axially mounted at a depth of 0.42 meters (1.38 feet) below IG-3. The second set of hoop-mounted gages (IG-5) occurs at a depth of 0.47 meters (1.54 feet) below IG-4. The last set (IG-6) is axially mounted at a depth of 0.6 meters (1.99 feet) below IG-5. The third illustration in this figure is a cross section of the instrumented FRP production pile showing a pair of vibrating wire strain gages, each mounted 58 millimeters (2.28 inches) in from the FRP outer wall and on opposite sides of the cylindrical pile.

Figure 208. Illustration. Layout of external durability instrumentation used in the FRP composite production pile. This figure presents two views showing the locations of the external bonded vibrating wire strain gages bonded to the FRP composite pile of pier 11. The first illustration represents a view toward the east showing a longitudinal section of the pile cap where it is attached to pier 11. The thickness of the pile cap is 1.22 meters (4.0 feet) and the diameter of the FRP pile is 0.62 meters (2.03 feet). The illustration shows that the upper 0.35 meters (1.15 feet) of the FRP pile is buried in the pile cap. The first vibrating wire strain gage (EG-1) is mounted longitudinally at 0.15 meters (0.49 feet) below the intersection of the FRP pile and pile cap. Gage EG-2 is also mounted longitudinally at 0.64 meters (2.1 feet) below this intersection. Gage EG-3 is mounted horizontally at 0.72 meters (2.36 feet) below the intersection of the pile and pile cap. Gage EG-4 is mounted longitudinally at a distance of 0.23 meters (0.75 feet) below EG-2, or 0.87 meters (2.85 feet) below the pile-cap intersection. Viewed toward the north, the illustration of this assembly is a cross section through the pile and pile cap. The drawing shows the same externally mounted gages, and their locations with respect to the pile-cap intersection and each other. The width of the pile cap is shown as 1.27 meters (4.17 feet).

Figure 209. Graph. Simplified stratigraphy along the Route 351 Bridge alignment. This figure shows the generalized stratigraphy of the soil column with respect to the pier alignment of the Route 351 Bridge and the Hampton River. The 12 piers and 2 end abutments are shown near the surface, from west on the left to east on the right. Five simplified soil layers, from 5 meters (16.4 feet) elevation to a depth of about 35 meters (115 feet), are described. The uppermost layer, between the near-surface and a depth of about 15 meters (49.2 feet) in the area below the deepest part of the river, is shown as a very soft silty clay with shell fragments and a median SPT value of 0. On the west side of the bridge below abutment A, and on the east side of the bridge below piers 8 through 12 and below abutment B, at a depth between the surface and about 10 meters (32.8 feet), is a layer of loose-to-medium silty sandy with occasional gravel and a median SPT value of 5. Below these layers, between a depth of about 10 to 20 meters (32.8 to 65.6 feet), is a layer of loose-to-medium silty sand with silty clay interbeds and a median SPT value of 11. The soil layer at a depth between about 20 and 27 meters (65.6 and 88.6 feet) is characterized as medium-dense silty sand with shell fragments and a median SPT value of 19. Below about 27 meters (88.6 feet) to about 33 meters (108 feet) is a layer of dense to very dense silty sand with shell fragments and a median SPT value of 51.

Figure 210. Graph. Simplified stratigraphy in the vicinity of the instrumented prestressed concrete pile installed at Pier 10. This figure consists of a simplified soil column identifying the general subsurface soil layers in the area surrounding pier 10. The soil stratigraphy is positioned alongside a point graph of the SPT N-values versus depth. The soil column depicts five distinct layers. Between the mudline and 5 meters (16.4 feet) depth, the soil is characterized as clay, silty, very soft, gray in color, and with shell fragments. Between a depth of 5 and 10 meters (16.4 and 32.8 feet), the soil is sand, silty, loose to medium dense, and green in color. The next layer, between 10 meters (32.8 feet) and about 19 meters (62.3 feet), the stratigraphy shows a deep layer of sand, silty, loose to medium dense, gray in color, and with shell fragments. A thin layer of gray sandy stiff clay is shown between about 19 and 20 meters (62.3 and 65.6 feet). Below 20 meters (65.6 feet) to the end of the borehole at 31.9 meters (105 feet), the soil is characterized as sand, silty, medium dense to dense, gray in color, and with shell fragments. The bottom of the prestressed concrete pile, shown alongside the generalized soil stratigraphy, is embedded in the upper clay and the following two sand layers. The scatter chart of SPT field N-values suggests that, within a discrete soil layer, the N-values increase with depth within that layer. Overall, however, the largest N-values are associated with the deepest measurements.

Figure 211. Graph. Simplified stratigraphy in the vicinity of the instrumented FRP composite pile installed at Pier 11. This figure consists of a simplified soil column identifying the general subsurface soil layers in the area surrounding pier 11. The soil stratigraphy is positioned alongside a point graph of the SPT N-values versus depth. The soil column depicts six distinct layers. Between the mudline and about 2.5 meters (8.2 feet) depth, the soil is characterized as clay, silty, sandy, soft, gray in color, and with shell fragments. Between a depth of 2.5 and 7.5 meters (8.2 and 24.6 feet), the soil is sand, silty, loose to medium dense, brown in color, and with shell fragments. For the next layer, between 7.5 meters (24.6 feet) and about 12 meters (39.4 feet), the stratigraphy shows a layer of sand, silty, medium dense, gray in color, and with shell fragments. Next is a thick layer of stiff clay, silty, sandy, with low plasticity, gray in color, and with shell fragments, occurring between 12 meters (39.4 feet) and about 20 meters (65.6 feet) in depth. Between 20 and 28 meters (65.6 and 91.9 feet), the soil is a thick layer of sand, silty, clayey, medium dense, gray in color, and with shell fragments. A thin layer of silty medium dense sand is shown between 28 and 30 meters (91.9 and 98.4 feet) depth. The bottom of the instrumented FRP composite pile, shown alongside the generalized soil stratigraphy, is embedded in the upper four clay and sand layers. The scatter chart of SPT field N-values suggests a very generalized trend of increasing N-values with depth in the soil column.

Figure 212. Graph. Driving records for instrumented production piles. This figure presents the pile driving records for the prestressed concrete and the FRP composite instrumented production piles. The X-axis is number of pile blows (0 to 100) per 305 millimeters (12 inches). The Y-axis is elevation in meters (positive 5 to negative 20) (positive 16.4 to negative 65.6 feet). The chart indicates that approximately 50 blows per 305 millimeters (12 inches) are required to drive the FRP composite pile at a depth of just under 20 meters (65.6 feet) below the surface mudline, while about 93 blows per 305 millimeters (12 inches) are needed to drive the prestressed concrete pile at approximately the same depth.

Figure 213. Graph. PDA recordings during restrike (Spiro and Pais 2002B). This figure consists of two graphs presenting force and velocity monitoring data obtained from the pile driving analyzer during restrike of the prestressed concrete pile (A) and the FRP composite pile (B). On each graph, the X-axis is time in milliseconds (0 to 50) and the Y-axis is force in kilonewtons (negative 4000 to positive 8000). A comparison of the two graphs shows that both types of piles exhibit similar dynamic behavior during driving. A force peak is found at time 0, which is the left end of a horizontal bar, superimposed on each graph, representing wave travel time for two pile lengths, or 2L over C, where L represents the pile length and C is the wave speed. Both graphs (A) and (B) show a large increase in the velocity record and a corresponding low force at about 2L over C. Between time zero and 0.5L over C, the data traces for the force and velocity records are very similar for both piles. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 214. Graph. PIT sounding on the prestressed concrete production pile before installation. This figure presents the PIT wave velocity curve for the prestressed concrete pile before driving. The X-axis is distance in feet (0 to 70) (0 to 21.3 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the graph indicates that the data were recorded on June 13, 2002; that wave speed was 13,065 feet per second (3,982 meters per second); and that the exact length of the prestressed concrete pile is 72.18 feet (22 meters). The graph also includes a rectangular representation of the production pile above the X-axis between 0 and about 72 feet (21.9 meters), corresponding to the length of the pile. The sounding data show only a small peak at time zero and minor toe reflection at the end of the data tracing.

Figure 215. Graph. PIT sounding on the prestressed concrete production pile after installation and restrike. This figure presents the PIT wave velocity curve for the prestressed concrete pile after driving. The X-axis is distance in feet (0 to 70) (0 to 21.3 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the graph indicates that the data were recorded on June 18, 2002; that wave speed was 13,200 feet per second (4,023 meters per second); and that the exact length of the prestressed concrete pile is 72.18 feet (22 meters). The graph also includes a rectangular representation of the production pile above the X-axis between 0 and about 72 feet (21.9 meters), corresponding to the length of the pile. The sounding data show only a small peak at time zero. The toe reflection peak at the end of the data tracing is barely detectable.

Figure 216. Graph. PIT sounding on the FRP composite production pile before installation. This figure presents the PIT wave velocity curve for the FRP composite pile before driving. The X-axis is distance in feet (0 to 70) (0 to 21.3 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the graph indicates that the data were recorded on March 5, 2002; that wave speed was 13,335 feet per second (4,064 meters per second); and that the exact length of the FRP composite pile is 69.6 feet (21.2 meters). The graph also includes a rectangular representation of the production pile above the X-axis between 0 and about 72 feet (21.9 meters), corresponding to the length of the pile. The sounding data show only a small peak at time zero and minor toe reflection at the end of the data tracing.

Figure 217. Graph. PIT sounding on the FRP composite production pile before installation. This figure presents the PIT wave velocity curve for the FRP composite pile after driving. The X-axis is distance in feet (0 to 70) (0 to 21.3 meters) and the Y-axis is velocity in inches per second (negative 0.4 to positive 1) (negative 10.2 millimeters to positive 25.4 millimeters). Text on the graph indicates that the data were recorded on June 18, 2002; that wave speed was 13,335 feet per second (4,064 meters per second); and that the exact length of the FRP composite pile is 69.6 feet (21.2 meters). The graph also includes a rectangular representation of the production pile above the X-axis between 0 and about 72 feet (21.9 meters), corresponding to the length of the pile. The sounding data show only a small peak at time zero.

Figure 218. Photo. Route 351 Bridge under construction on December 30, 2002. This figure is a color photograph of the Route 351 bridge taken during its construction. In the photograph, piers 7 through 11 are shown installed, with the bridge deck partially completed. On pier 11, the cylindrical instrumented FRP composite pile can clearly be seen. The instrumented prestressed concrete pile at pier 10 is partially visible. The photograph also shows the formwork on the bridge deck being prepared for cement pouring.

Figure 219. Graph. Load distributions on November 7, 2002, and December 30, 2002, for the prestressed concrete production pile at Pier 10. This figure presents the estimated load distributions calculated for the prestressed concrete production pile during November and December 2002. The X-axis is load in kilonewtons (0 to 250) and the Y-axis is depth in meters (0 to 25) (0 to 82 feet). The graph indicates that the load measured near the surface at the end of December is almost twice the load measured near the surface in early November. With increased depth, the difference between the two temporal measurements is not as great. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 220. Graph. Load distributions on November 7, 2002, and December 30, 2002, for the FRP composite production pile at Pier 11. This figure presents the estimated load distributions calculated for the FRP composite production pile during November and December 2002. The X-axis is load in kilonewtons (0 to 350) and the Y-axis is depth in meters (0 to 25) (0 to 82 feet). The load measured near the surface at the end of November is 150 kilonewtons and in late December the near-surface load increased to about 275 kilonewtons. With increased depth, the curves from the two temporal measurements converge. The conversion factor for kilonewtons is 1 kilonewton equals 225 poundforce.

Figure 221. Graphs. Interface shear test results, Density sand-to-Lancaster FRP shell interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement for the Density sand-to-Lancaster FRP shell interface (bottom). On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the upper graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-Lancaster FRP shell interface, the upper graph shows that the greatest peak interface shear stress is about 70 kilopascals, and occurs at a relative density of 65.7 percent and a constant normal pressure of 200 kilopascals. A peak interface shear stress of about 40 kilopascals is shown for a relative density of 62.4 percent and constant normal pressure of 113.2 kilopascals. Lowering the constant normal pressure to 63.2 kilopascals with a relative density of 64.3 percent, the peak interface shear stress is about 22 kilopascals. The lowest peak shear stress is shown for a relative sand density of 57.5 percent and a constant normal pressure of 38.2 kilopascals. The graph on the bottom shows that vertical displacement is minor (plus or minus about 0.02 millimeters (7.9 times 10 to the negative 4 power inches)) for a combination of either the highest constant normal stress and the greatest relative density or the lowest constant normal stress and the smallest relative density shown in the top graph. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 222. Graphs. Interface shear test results, Density sand-to-Hardcore FRP shell interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Density sand-to-Hardcore FRP shell interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-Hardcore FRP shell interface, the top graph shows that the greatest peak interface shear stress is about 110 kilopascals, and occurs at a relative density of 68.3 percent and a constant normal pressure of 200 kilopascals. A peak interface shear stress of about 55 kilopascals is shown for a relative density of 66.3 percent and constant normal pressure of 100 kilopascals. Lowering the constant normal pressure to 50 kilopascals with a relative density of 66 percent, the peak interface shear stress is about 28 kilopascals. The lowest peak shear stress is shown for a relative sand density of 61.4 percent and a constant normal pressure of 22.5 kilopascals. The bottom graph shows that vertical displacement is positive for all combinations of relative density and constant normal stress presented in the top graph. The graph shows that the largest vertical displacement occurs for the 66.3-percent relative density and 100-kilopascal constant normal stress. The smallest vertical displacement is shown for the combination of a relative sand density of 61.4 percent and a constant normal stress of 22.5 kilopascals. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 223. Graphs. Interface shear test results, Density sand-to-Hardcore FRP plate (untreated) interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Density sand and the untreated Hardcore FRP plate interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-Hardcore FRP plate interface, the top graph shows that the greatest peak interface shear stress is about 110 kilopascals, and occurs at a relative density of 66.2 percent and a constant normal pressure of 199.73 kilopascals. A peak interface shear stress of about 60 kilopascals is shown for a relative density of 63.4 percent and constant normal pressure of 107.3 kilopascals. Lowering the constant normal pressure to 50.06 kilopascals with a relative density of 61.5 percent, the peak interface shear stress is about 30 kilopascals. The lowest peak shear stress is shown for a relative sand density of 64.6 percent and a constant normal pressure of 26.15 kilopascals. The bottom graph shows that vertical displacement is near zero or negative for all combinations of relative density and constant normal stress presented in the top graph. The graph shows that negative vertical displacements occur for the 63.4-percent relative density and 107.3-kilopascal constant normal stress, and for the 66.2-percent relative density and 199.73-kilopascal constant normal stress. Near-zero vertical displacement is shown for the combination of a relative sand density of 61.5 percent and a constant normal stress of 50.06 kilopascals, and for the combination of a relative sand density of 64.6 percent and a constant normal stress of 26.15 kilopascals. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 224. Graphs. Interface shear test results, Density sand-to-treated Hardcore FRP plate interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Density sand and the treated FRP plate interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-treated FRP plate interface, the top graph shows that the greatest peak interface shear stress is about 115 kilopascals, and occurs at a relative density of 64.9 percent and a constant normal pressure of 183.5 kilopascals. A peak interface shear stress of about 62 kilopascals is shown for a relative density of 63.7 percent and constant normal pressure of 107.17 kilopascals. Lowering the constant normal pressure to 50.06 kilopascals at a relative density of 64.7 percent, the peak interface shear stress is about 30 kilopascals. The lowest peak shear stress is shown for a relative sand density of 62.6 percent and a constant normal pressure of 26.15 kilopascals. The bottom graph shows that vertical displacement for all density-shear stress combinations shown in the top graph is initially negative for 1 to 2 millimeters (0.04 to 0.08 inches) of interface displacement, and then increases to a range between near zero and positive 0.16 millimeters (0.0063 inches). The greatest vertical displacement, 0.16 millimeters (0.0063 inches), is shown for the combination of a 64.9-percent relative density and a 183.5-kilopascal constant normal pressure. The smallest vertical displacement, near zero, occurs for the combination of a relative density of 63.7 percent and a constant normal pressure of 107.17 kilopascals. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 225. Graphs. Interface shear test results, Density sand-to-plastic interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Density sand and the plastic interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-plastic interface, the top graph shows that the greatest peak interface shear stress is about 105 kilopascals, and occurs at a relative density of 61.1 percent and a constant normal pressure of 199.72 kilopascals. A peak interface shear stress of about 92 kilopascals is shown for a relative density of 61.9 percent and constant normal pressure of 180.37 kilopascals. Lowering the constant normal pressure to 105.37 kilopascals at a relative density of 60.8 percent, the peak interface shear stress is about 55 kilopascals. At a constant normal pressure of 56.77 kilopascals and a relative sand density of 61.1 percent, the interface shear stress is about 32 kilopascals. The lowest peak shear stress is shown for a relative sand density of 65.9 percent and a constant normal pressure of 29.11 kilopascals. The bottom graph shows that vertical displacement for density-shear stress combinations shown in the top graph is slightly negative or near zero for the three highest constant normal pressures of 199.72, 180.3, and 105.37 kilopascals, and slightly positive for the two lowest constant normal pressures of 56.77 and 29.11 kilopascals. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 226. Graphs. Interface shear test results, Density sand-to-concrete interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Density sand and the concrete interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-concrete interface, the top graph shows that the greatest peak interface shear stress is about 112 kilopascals, and occurs at a relative density of 64 percent and a constant normal pressure of 175.4 kilopascals. A peak interface shear stress of about 68 kilopascals is shown for a relative density of 63.3 percent and constant normal pressure of 101.9 kilopascals. Lowering the constant normal pressure to 51.5 kilopascals at a relative density of 61.9 percent, the peak interface shear stress is about 35 kilopascals. The lowest peak interface shear stress, about 18 kilopascals, is shown for a relative sand density of 63.2 percent and a constant normal pressure of 23.7 kilopascals. The bottom graph shows that, after an interface displacement of about 3 to 4 millimeters (0.118 to 0.157 inches), the vertical displacement for all sand relative density-constant normal pressure combinations shown in the top graph is narrowly positive 0.1 millimeters (0.004 inches). The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 227. Graphs. Interface shear test results, Density sand-to-steel interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Density sand and the steel interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Density sand-to-steel interface, the top graph shows that the greatest peak interface shear stress is about 105 kilopascals, and occurs at a relative density of 65.4 percent and a constant normal pressure of 202.33 kilopascals. A peak interface shear stress of about 60 kilopascals is shown for a relative density of 64.1 percent and constant normal pressure of 101.85 kilopascals. Lowering the constant normal pressure to 51.27 kilopascals at a relative density of 65.9 percent, the peak interface shear stress is about 30 kilopascals. The lowest peak interface shear stress, about 12 kilopascals, is shown for a relative sand density of 67 percent and a constant normal pressure of 25.57 kilopascals. The bottom graph shows that the vertical displacement for the 65.4-percent sand relative density and the 202.33-kilopascal constant normal pressure combination is about negative 0.05 millimeters (negative 0.002 inches). The remaining three combinations of pressure and density shown in the top graph are all near zero or slightly positive to less than 0.05 millimeters (0.002 inches). The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 228. Graphs. Interface shear test results, Model sand-to-Lancaster FRP shell interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the FRP shell interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-FRP shell interface, the top graph shows that the greatest peak interface shear stress is about 105 kilopascals, and occurs at a relative density of 64.9 percent and a constant normal pressure of 200 kilopascals. A peak interface shear stress of about 50 kilopascals is shown for a relative density of 63.2 percent and constant normal pressure of 100 kilopascals. Lowering the constant normal pressure to 50 kilopascals at a relative density of 58.7 percent, the peak interface shear stress is about 25 kilopascals. The lowest peak interface shear stress, about 10 kilopascals, is shown for a relative sand density of 57.8 percent and a constant normal pressure of 22.5 kilopascals. The bottom graph shows that the vertical displacement for the 64.9-percent sand relative density and the 200-kilopascal constant normal pressure combination is about slightly negative to near zero. The remaining three combinations of pressure and density shown in the top graph all show vertical displacements above zero between 0.05 and 0.1 millimeters (0.002 and 0.004 inches). The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 229. Graphs. Interface shear test results, Model sand-to-Hardcore FRP shell interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the Hardcore FRP shell interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-Hardcore FRP shell interface, the top graph shows that the greatest peak interface shear stress is about 112 kilopascals, and occurs at a relative density of 67.9 percent and a constant normal pressure of 200 kilopascals. A peak interface shear stress of about 55 kilopascals is shown for a relative density of 60.9 percent and constant normal pressure of 100 kilopascals. Lowering the constant normal pressure to 50 kilopascals at a relative density of 64.4 percent, the peak interface shear stress is about 28 kilopascals. The lowest peak interface shear stress, about 15 kilopascals, is shown for a relative sand density of 56.1 percent and a constant normal pressure of 22.5 kilopascals. The bottom graph shows that the vertical displacement for all combinations of relative sand density and constant normal pressure is in the positive range. The smallest vertical displacement is associated with a relative sand density of 60.9 percent and a constant normal pressure of 100 kilopascals. The largest vertical displacement, about positive 0.15 millimeters (0.006 inches), is for a relative sand density of 56.1 percent and a constant normal pressure of 22.5 kilopascals. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 230. Graphs. Interface shear test results, Model sand-to-Hardcore FRP plate (untreated) interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the untreated FRP plate interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-untreated FRP plate interface, the top graph shows that the greatest peak interface shear stress is about 112 kilopascals, and occurs at a relative density of 66.4 percent and a constant normal pressure of 183.5 kilopascals. A peak interface shear stress of about 65 kilopascals is shown for a relative density of 64.7 percent and constant normal pressure of 107.17 kilopascals. Lowering the constant normal pressure to 50.06 kilopascals at a relative density of 59.1 percent, the peak interface shear stress is about 32 kilopascals. The lowest peak interface shear stress, about 18 kilopascals, is shown for a relative sand density of 65.4 percent and a constant normal pressure of 26.15 kilopascals. The bottom graph shows that vertical displacement for all density-interface shear stress combinations shown in the top graph is initially negative for about 1 millimeter (0.04 inches) of interface displacement, and then increases to a range between negative 0.06 millimeters (negative 0.0024 inches) and positive 0.08 millimeters (positive 0.003 inches). The largest vertical displacement, positive 0.08 millimeters (positive 0.003 inches), occurs for the combination of a relative density of 65.4 percent and a constant normal pressure of 26.15 kilopascals. A vertical displacement of negative 0.06 millimeters (negative 0.0024 inches) is shown for the combination of a 64.7-percent relative density and a 107.17-kilopascal constant normal pressure, and for a 59.1-percent relative sand density and a 50.06-kilopascal constant normal pressure. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 231. Graphs. Interface shear test results, Model sand-to-treated Hardcore FRP plate interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the treated FRP plate interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-treated FRP plate interface, the top graph shows that the greatest peak interface shear stress is about 118 kilopascals, and occurs at a relative density of 62 percent and a constant normal pressure of 152.2 kilopascals. A peak interface shear stress of about 80 kilopascals is shown for a relative density of 60.1 percent and constant normal pressure of 107.3 kilopascals. Lowering the constant normal pressure to 50.08 kilopascals at a relative density of 65 percent, the peak interface shear stress is about 40 kilopascals. The lowest peak interface shear stress, about 22 kilopascals, is shown for a relative sand density of 63 percent and a constant normal pressure of 26.15 kilopascals. The bottom graph shows that vertical displacement for all density-interface shear stress combinations shown in the top graph is initially negative for about 1 to 2 millimeters (0.04 to 0.08 inches) of interface displacement, and then increases rapidly to a range between positive 0.1 millimeters (positive 0.004 inches) and positive 0.26 millimeters (positive 0.01 inches). The largest vertical displacement, positive 0.26 millimeters (positive 0.01 inches), occurs for the combination of a relative density of 65 percent and a constant normal pressure of 50.08 kilopascals. A vertical displacement of 0.18 millimeters (0.0071 inches) is shown for the combination of a 63-percent relative density and a 26.15-kilopascal constant normal pressure. Similar vertical displacements of about positive 0.1 millimeters (positive 0.004 inches) are shown for the combinations of 62 percent sand density and 152.2 kilopascals constant normal pressure, and 60.1 percent sand density and 107.3 kilopascals constant normal pressure. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 232. Graphs. Interface shear test results, Model sand-to-plastic interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the treated FRP plate interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-plastic interface, the top graph shows that the greatest peak interface shear stress is about 128 kilopascals, and occurs at a relative density of 62.4 percent and a constant normal pressure of 199.7 kilopascals. A peak interface shear stress of about 105 kilopascals is shown for a relative density of 66.4 percent and constant normal pressure of 152.2 kilopascals. Lowering the constant normal pressure to 107.25 kilopascals at a relative density of 65.1 percent, the peak interface shear stress is about 75 kilopascals. At a relative sand density of 58.1 percent and a constant normal pressure of 50.1 kilopascals, the interface shear stress is about 28 kilopascals. The lowest peak interface shear stress, about 15 kilopascals, is shown for a relative sand density of 64.4 percent and a constant normal pressure of 26.15 kilopascals. The bottom graph shows that vertical displacements for all sand density-interface shear stress combinations shown in the top graph, except the highest constant normal pressure of 199.7 kilopascals with a sand density of 62.4 percent, are similar in a range of about positive 0.05 to 0.08 millimeters (positive 0.002 to 0.003 inches). A vertical displacement of about negative 0.1 millimeters (negative 0.004 inches) is shown for a combination of 62.4 percent sand density and 199.7 kilopascals constant normal pressure. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 233. Graphs. Interface shear test results, Model sand-to-concrete interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the concrete interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-concrete interface, the top graph shows that the greatest peak interface shear stress is about 128 kilopascals, and occurs at a relative density of 68.2 percent and a constant normal pressure of 175.4 kilopascals. A peak interface shear stress of about 72 kilopascals is shown for a relative density of 61.9 percent and constant normal pressure of 101.9 kilopascals. Lowering the constant normal pressure to 51.5 kilopascals at a relative density of 58.7 percent, the peak interface shear stress is about 38 kilopascals. The lowest peak interface shear stress, about 20 kilopascals, is shown for a relative sand density of 58.9 percent and a constant normal pressure of 23.7 kilopascals. The bottom graph shows that vertical displacements for all sand density-interface shear stress combinations shown in the top graph, except the lowest constant normal pressure of 23.7 kilopascals with a sand density of 58.9 percent, are similar at about positive 0.1 millimeters (positive 0.004 inches). A vertical displacement of about 0.2 millimeters (negative 0.008 inches) is shown for a combination of 58.9 percent sand density and 23.7 kilopascals constant normal pressure. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 234. Graphs. Interface shear test results, Model sand-to-steel interface. This figure consists of two graphs that present the results of the interface shear test (top) and the vertical displacement (bottom) for the Model sand and the steel interface. On both graphs, the X-axis is interface displacement in millimeters (0 to 15) (0 to 0.59 inches). The Y-axis of the top graph is interface shear stress in kilopascals (0 to 150). The Y-axis of the bottom graph is vertical displacement in millimeters (negative 0.5 to positive 0.5) (negative 0.02 to positive 0.02 inches). The obvious trend shown is an increase in interface shear stress with an increase in constant normal pressure. For the Model sand-to-steel interface, the top graph shows that the greatest peak interface shear stress is about 122 kilopascals, and occurs at a relative density of 67 percent and a constant normal pressure of 202.33 kilopascals. A peak interface shear stress of about 62 kilopascals is shown for a relative density of 64.3 percent and constant normal pressure of 101.85 kilopascals. Lowering the constant normal pressure to 51.27 kilopascals at a relative density of 62.5 percent, the peak interface shear stress is about 31 kilopascals. The lowest peak interface shear stress, about 18 kilopascals, is shown for a relative sand density of 63.5 percent and a constant normal pressure of 26.08 kilopascals. The bottom graph shows that vertical displacements for all sand density-interface shear stress combinations shown in the top graph, except the lowest constant normal pressure of 26.08 kilopascals with a sand density of 63.5 percent, are similar in a narrow range of about negative 0.02 to 0.04 millimeters (negative 0.0008 to 0.0016 inches). A vertical displacement of about positive 0.1 millimeters (positive 0.004 inches) is shown for a combination of 63.5 percent sand density and 26.08 kilopascals constant normal pressure. The conversion factor for kilopascals is 1 kilopascal equals 0.145 poundforce per square inch.

Figure 235. Graph. FRP moisture concentration profile, inner radius dry, outer radius saturated. This figure compares the temporal distribution of moisture content of the FRP composite pile assuming the outer radius is saturated and the inner radius is dry. The X-axis is a formula relating the inner and outer radii of the FRP tube on a scale of negative 0.5 to positive 0.5. The formula is the quotient of the sum of the product of 2 times R minus the sum of R subscript O plus R subscript I divided by 2 times the sum of R subscript O minus R subscript I. The Y-axis is the concentration of moisture in percent (0 to 0.5). The overall curve for 100 days is approximately linear. For periods of 30 days, 10 days, and 1 day, the curve is increasingly exponential. Each of the four curves begins at a moisture concentration of 0.00 percent at an X-axis reading of negative 0.5 and rises to a moisture concentration of approximately 0.425 percent at an X-axis reading of positive 0.5.

Figure 236. Graph. FRP moisture concentration profile, inner and outer radii saturated. This figure compares the temporal distribution of moisture content of the FRP composite pile assuming both the inner and outer radii of the FRP composite tube are saturated. The X-axis is a formula relating the inner and outer radii of the FRP tube on a scale of negative 0.5 to positive 0.5. The formula is the quotient of the sum of the product of 2 times R minus the sum of R subscript O plus R subscript I divided by 2 times the sum of R subscript O minus R subscript I. The Y-axis is the concentration of moisture in percent (0 to 0.5). Curves for 100 days, 30 days, 10 days, and 1 day are given. All four curves begin at a moisture concentration of approximately 0.425 percent at an X-axis reading of negative 0.5, descend as the X-axis reading approaches 0, and then rise to a moisture concentration of approximately 0.425 at an X-axis reading of positive 0.5. The 100-day curve descends the least, and the 1-day curve descends the most.

Figure 237. Photo. VTRC pushout test setup. This figure is a color photograph that shows the setup used by the Virginia Transportation Research Council to conduct the pushout tests on the Lancaster composite piles. The photograph shows a slice of the test pile on a block with a hole, which allows the block to support the FRP shell but permits the concrete core to pass through.

Figure 238. Photo. VT pushout test setup. This figure is a color photograph of the setup used by Virginia Tech to perform the pushout tests on the Lancaster composite piles. This photograph shows a section of steel pipe, within which the FRP shell is bonded. This allows the concrete core to be pushed out of the FRP shell.

Figure 239. Photo. Creep bending test setup. This figure is a color photograph of the four-point creep bending test apparatus used to conduct the creep bending test. The apparatus is set up inside a tent. The photograph shows a horizontally supported 5.5-meter (18-foot) section of the 610-millimeter (24-inch) diameter Lancaster composite pile. Two large 7.2-kip dead weights are shown suspended from the supporting beam.

Figure 240. Graph. Creep deflection test results. This figure illustrates the results of the creep deflection test conducted on the Lancaster composite pile. The X-axis is elapsed time in hours (0 to 10,000) and the Y-axis is centerline deflection in millimeters (0 to 3) (0 to 0.118 inches). The instantaneous centerline deflection response near time zero is just beyond 1 millimeter (0.04 inches). Measurements of centerline deflection are plotted on the chart and show a steady increase between the instantaneous response at about 1 millimeter and the last measurement of about 2.2 millimeters (0.086 inches) after 377 days.

Figure 241. Map. Location of field tests. This figure shows the locations of the geotechnical field investigations completed at the Route 351 bridge test site. Text indicates that the map, which is shown at a scale of 1 to 250, was adapted from a copy of Sheet 56 of the Project Road Plans. On the left side of the map, which is the west direction, is the Hampton River running in the north-south direction. The pile load test site is shown on a spit of land that parallels the east side of the river. The test pile layout illustrates the locations of six test sites along a north-south axis. The northern-most site it SPT-1, followed by CPT034. Below CPT034 is the location of the pile load test and CPT-1. Below this boring site is SPT-2, followed by CPT033, and CPT-2.

Figure 242. Chart. Boring B-1 (SPT-1). This figure is a two-page test boring record for SPT-1, and presents the soil stratigraphy and blow record chart. The X-axis along the chart is number of blows per foot (0 to 50) and adjacent to that is number of blows per 6 inches (15.2 cm). The Y-axis is depth in feet (0 to 100.5) (0 to 30.6 meters). The soil description reveals a 1-inch (25.4-millimeter) layer of topsoil at the surface. Below the topsoil to a depth of 2 feet (0.61 meters) is a layer of fill, described as brown silty fine sand; traces of clay, shell fragments, and gravel; moist, medium compact, SM. Between 2 and 4 feet (0.61 and 1.22 meters) is a layer identified as possible fill, described as brown, silty clayey fine sand; traces of shell fragments and gravel; moist, loose, SC. Between 4 and 42 feet (1.22 and 12.8 meters), the soil description has minor variations but is basically as follows: brown silty fine sand; shell fragments and traces of clay; wet, medium compact, SM. Between 42 and 47 feet (12.8 and 14.3 meters), the soil is gray silty fine sandy clay; traces of shell fragments; wet, stiff, CL. Between 47 and 52 feet (14.3 and 15.8 meters), the soil layer is gray silty clayey fine sand; traces of shell fragments; wet, medium compact, SC. Between 52 and 62 feet (15.8 and 18.9 meters), the soil is gray silty fine sand; traces of clay and shell fragments; wet, medium compact, SM. Between 62 and 67 feet (18.9 and 20.4 meters), the soil layer is described as gray clayey silty fine sand; traces of shell fragments; wet, medium compact, SM-SC. From 67 to 77 feet (20.4 and 23.5 meters), the soil is gray clayey silty fine sand; traces of shell fragments; wet, medium compact, SC. Between 77 feet (23.5 meters) and the end of the boring at 100.5 feet (30.6 meters), the soil is described as gray silty fine sand; traces of clay and shell fragments; wet, medium compact, SM. Adjacent to the soil stratigraphy is the blow record. Within the upper 25 feet (7.6 meters), the blow record indicates that blows generally range from 5 to 15 per foot (0.3 meters). Between 25 and 45 feet (7.6 and 13.7 meters), the blow record indicates that the range is between about 5 and 25 blows per foot (0.3 meters). From a depth of 50 feet (15.2 meters) to the end of the boring at 100.5 feet (30.6 meters), the number of blows per foot varies between 15 and 30.

Figure 243. Chart. Boring B-2 (SPT-2). This figure is a two-page test boring record for SPT-2, and presents the soil stratigraphy and blow record chart. The X-axis along the chart is number of blows per foot (0 to 50) and adjacent to that is number of blows per 6 inches (15.2 cm). The Y-axis is depth in feet (0 to 100.5) (0 to 30.6 meters). The soil description reveals a 1-inch (25.4-millimeter) layer of topsoil at the surface. Below the topsoil to a depth of 2 feet (0.61 meters) is a layer of possible fill, described as brown silty fine sand; organics, traces of clay, and shell fragments; moist, loose, SM. Between 2 and 6 feet (0.61 and 1.83 meters) is another layer identified as possible fill, described as brown silty fine sandy clay; interspersed shell fragments; wet, soft, CL. Between 6 and 52 feet (1.83 and 15.8 meters), the soil description has minor variations but is basically as follows: brown silty fine sand; shell fragments and traces of clay; wet, very loose, SM. Between 52 and 62 feet (15.8 and 18.9 meters), the soil is gray clayey silty fine sand; traces of shell fragments; wet, medium compact, SM-SC. Between 62 and 72 feet (18.9 and 21.9 meters), the soil layer is gray silty clayey fine sand; traces of shell fragments; wet, medium compact, SC. Between 72 and 92 feet (21.9 and 28 meters), the soil is gray clayey silty fine sand; traces of shell fragments; wet, medium compact, SM-SC. Between 92 and 97 feet (28 and 29.6 meters), the soil layer is described as gray silty clayey fine sand; traces of shell fragments; wet, medium compact, SC. From 97 feet (29.6 meters) to the end of the boring at 100.5 feet (30.6 meters), the soil is gray clayey silty fine sand; traces of shell fragments; wet, medium compact, SM-SC. Adjacent to the soil stratigraphy is the blow record. Within the upper 30 feet (9.1 meters), the blow record indicates that blows generally range from 5 to 15 per foot (0.3 meters). Between 30 and 50 feet (9.1 and 15.2 meters), the blow record indicates that the range is between about 15 and 30 blows per foot (0.3 meters). From a depth of 50 feet (15.2 meters) to the end of the boring at 100.5 feet (30.6 meters), the number of blows per foot is generally consistent and varies narrowly between 20 and 25.

Figure 244. Chart. CPT-1. This figure consists of four charts that present the results obtained from the electric cone penetrometer test at CPT-1. These results include pore pressure, the friction ratio, the cone bearing, and the sleeve stress. The first chart presents pore pressure. The upper X-axis is enlarged-scale pore pressure in meters (0 to 40) (0 to 131.2 feet) and the lower X-axis is full-scale pore pressure in meters (0 to 400) (0 to 1,312 feet). The Y-axis is depth in meters (0 to 30) (0 to 98.4 feet). The data show that the pore pressure on the enlarged scale, between a depth of about 12 meters (39.4 feet) and 30 meters (98.4 feet), varies widely between 0 and 40 meters (0 to 131.2 feet). On the full-scale chart, between the same depths, the pore pressure varies irregularly between 0 and 400 meters (0 to 1,312 feet). The next data chart presents the friction ratio. The X-axis is the ratio in percent (0 to 8). The Y-axis is depth in meters (0 to 30) (0 to 98.4 feet). The data indicate that the friction ratio varies mostly between 0 and 2 percent. The next chart includes data for the cone bearing. The upper X-axis is the enlarged scale for the cone bearing in bars (0 to 40) and the lower X-axis is the full scale for the cone bearing in bars (0 to 400). The cone bearing appears to increase with depth. The last chart presents a depth profile for the sleeve stress, indicated on the X-axis in bars (0 to 3). The sleeve stress appears to increase very gradually with depth.

Figure 245. Chart. CPT-2. This figure consists of four charts that present the results obtained from the electric cone penetrometer test at CPT-2. These results include pore pressure, the friction ratio, the cone bearing, and the sleeve stress. The first chart presents pore pressure. The upper X-axis is enlarged-scale pore pressure in meters (0 to 40) (0 to 131.2 feet) and the lower X-axis is full-scale pore pressure in meters (0 to 400) (0 to 1,312 feet). The Y-axis is depth in meters (0 to 30) (0 to 98.4 feet). The data show that the pore pressure on the enlarged scale, between a depth of about 12 meters (39.4 feet) and 30 meters (98.4 feet), varies widely between 0 and 40 meters (0 to 131.2 feet). On the full-scale chart, between the same depths, the pore pressure varies irregularly between 0 and 400 meters (0 to 1,312 feet). The next data chart presents the friction ratio. The X-axis is the ratio in percent (0 to 8). The Y-axis is depth in meters (0 to 30) (0 to 98.4 feet). The data indicate that the friction ratio varies mostly between 0 and 2 percent. The next chart includes data for the cone bearing. The upper X-axis is the enlarged scale for the cone bearing in bars (0 to 40) and the lower X-axis is the full scale for the cone bearing in bars (0 to 400). The cone bearing appears to increase with depth. The last chart presents a depth profile for the sleeve stress, indicated on the X-axis in bars (0 to 3). The sleeve stress appears to increase very gradually with depth.

Figure 246. Chart. CPT033. This figure is a four-page cone penetrometer test log for CPT033. It presents depth profiles for the friction, cone resistance, pore pressure, and soil behavior type. Text on the chart indicates that the cone size is 10 square centimeters (1.55 square inches); the hole was drilled by the Virginia Department of Transportation on August 16, 2001; groundwater occurs at 5 feet (1.5 meters); and the surface elevation is 5 feet (1.5 meters). The Y-axis for all the parameters is depth in feet (0 to 94) (0 to 28.6 meters) as well as elevation in feet (positive 5 to negative 90) (positive 1.5 to negative 27.4 meters). The first profile is for friction. The X-axis is friction measured in tons per square feet (0 to 1 TSF). Between the surface and about 30 feet (9.1 meters), the friction increases gradually to about 0.4 TSF. Between 30 and 40 feet (9.1 and 12.2 meters), the friction varies little between 0.2 and 0.5 TSF. Between 40 and 70 feet (12.2 and 21.3 meters), friction declines to 0.2 TSF or less. From 70 feet (21.3 meters) to the bottom of the drill hole at 95 feet (29 meters), the friction increases gradually to 0.1 TSF. The next profile is cone resistance measured in units of TSF (0 to 60). The data trace indicates that cone resistance between the surface and about 10 feet (3.05 meters) is below 5 TSF. From 10 feet (3.05 meters) to about 43 feet (13.1 meters), the cone resistance increases gradually to about 15 TSF. From 43 to 73 feet (13.1 to 22.2 meters), cone resistance decreases to about 5 TSF. Between 73 feet (13.1 meters) and the bottom of the drill hole at 95 feet (29 meters), cone resistance increases gradually to a maximum of 30 TSF. The next profile is for pore pressure, measured on the X-axis in units of TSF (0 to 0.8). Pore pressure is detectable between 5 and 8 feet (1.5 and 2.4 meters), and again at between 47 and 74 feet (14.3 and 22.5 meters), where it increases rapidly to a maximum of 0.8 TSF. Between 74 and 94 feet (22.5 and 28.6 meters), pore pressure decreases from the maximum, with values varying between 0.1 and 0.4 TSF. Alongside the parameter profiles is a chart describing soil behavior type. These charts indicate that the soil is primarily sandy silt to clayey silt or silty sand to sandy silt. A layer of clay is shown between 5 and 10 feet (1.5 and 3.05 meters). Below 70 feet (21.3 meters), the soil type is primarily silty sand to sandy silt and sand to silty sand. The conversion factor for tons per square foot is 1 ton per square foot equals 9.75 metric tons per square meter.

Figure 247. Chart. CPT034. This figure is a three-page cone penetrometer test log for CPT034. It presents depth profiles for the friction, cone resistance, pore pressure, and soil behavior type. Text on the chart indicates that the cone size is 10 square centimeters (1.55 square inches); the hole was drilled by the Virginia Department of Transportation on August 16, 2001; groundwater occurs at 5 feet (1.5 meters); and the surface elevation is 5 feet (1.5 meters). The Y-axis for all the parameters is depth in feet (0 to 85) (0 to 25.9 meters) as well as elevation in feet (positive 5 to negative 80) (positive 1.5 to negative 24.4 meters). The first profile is for friction. The X-axis is friction measured in tons per square feet (0 to 1 TSF). Between the surface and about 35 feet (10.7 meters), the friction is between 0.2 and 0.3 TSF. Between 35 and 40 feet (10.7 and 12.2 meters), there is a small increase in friction to about 0.65 TSF. Between 40 and 58 feet (12.2 and 17.7 meters), friction declines and varies between 0.2 and 0.6 TSF. From 58 feet (17.7 meters) to about 70 feet (21.3 meters), the friction is generally constant at about 0.2 TSF. Below 70 feet (21.3 meters) to the bottom of the drill hole at 85 feet (25.9 meters), the friction increases gradually to 0.8 TSF. The next profile is cone resistance measured in units of TSF (0 to 60). The data trace indicates that cone resistance between the surface and about 10 feet (3.05 meters) is mostly below 5 TSF. From 10 feet (3.05 meters) to about 36 feet (11 meters), the cone resistance increases gradually to about 10 TSF. From 36 to 42 feet (11 to 12.8 meters), cone resistance increases to about 15 TSF. Between 42 feet and 75 feet (12.8 and 22.9 meters), the profile indicates that cone resistance is between 5 and 10 TSF. From 75 feet (22.9 meters) to the bottom of the drill hole at 85 feet (25.9 meters), cone resistance increases gradually to a maximum of about 17 TSF. The next profile is for pore pressure, measured on the X-axis in units of TSF (0 to 0.8). Pore pressure is detectable between 44 and 85 feet (13.4 and 25.9 meters). At 44 feet (13.4 meters), it increases sharply to a maximum value of 0.8 TSF, and fluctuates widely between 0 and 0.8 TSF in the 44- to 85-foot (13.4- to 25.9-meter) depth interval. Alongside the parameter profiles is a chart describing soil behavior type. These charts indicate that the soil above 5 feet (1.5 meters) is primarily silty sand to sandy silt. A layer of clay is shown between 6 and 10 feet (1.83 and 3.05 meters). From 10 to about 15 feet (3.05 to 4.57 meters), the soil becomes primarily sand to silty sand. Between 15 and 22 feet (4.57 and 6.71 meters), the description is mostly silty sand to sandy silt with sandy silt to clayey silt also appearing between 22 and 30 feet (4.57 and 9.1 meters). From 30 to 44 feet (9.1 to 13.4 meters), the soil becomes mostly sand to silty sand, with some silty sand to sandy silt. Between 44 and 75 feet (13.4 and 22.9 meters), the soil composition is primarily sandy silt to clayey silt, with some clayey silt to silty clay. A narrow band of silty sand to sandy silt is shown between 72 and 75 feet (21.9 and 22.9 meters). From 75 feet (22.9 meters) to the end of the borehole at 85 feet (25.9 meters), the soil is mostly sand to silty sand, with a narrow layer of silty sand to sandy silt near the bottom. The conversion factor for tons per square foot is 1 ton per square foot equals 9.75 metric tons per square meter.

Figure 248. Chart. DMT. This figure consists of three charts that present the results of the dilatometer test. The first chart presents a depth profile of the logarithmic material index, shown on the X-axis (0.1 to 0.6, corresponding to clay and 1.2 to 10, corresponding to sand). The Y-axis is depth in meters (0 to 10) (0 to 32.8 feet). The chart initially shows data points moving from sand to clay between the surface and a depth of 3 meters (9.85 feet). Between 3 and 6 meters (9.85 and 19.7 feet), the data points are primarily in the sand region. The next chart presents the DMT readings for P subscript zero and for P subscript 1 in bars (0 to 40) on the X-axis. The Y-axis is depth in meters (0 to 10) (0 to 32.8 feet). The DMT readings overall are slightly higher for P subscript zero. Except for the upper 1 meter (3.3 feet), both readings are below 5 bars for the 1- to 6-meter (3.3- to 19.7-foot) depth interval. The last chart presents the depth profile for thrust. The X-axis is logarithmic kilogram-force (100 to 10,000) (980.6 to 98,060 newtons). The Y-axis is depth in meters (0 to 10) (0 to 32.8 feet). The chart indicates that, within the upper 1 meter (3.3 feet), the thrust is high, but gradually decreases to near 1000 with depth.

Equations

Equation 1. The radial confinement pressure, sigma subscript R, equals the axial strain of the concrete-filled FRP tube stub, epsilon subscript CC, times the following quotient: the Poisson ratio of concrete, upsilon subscript C, minus the Poisson ratio of the FRP shell, upsilon subscript FRP, divided by the following: the sum of the quotient of the radius of the FRP tube, R subscript FRP, divided by the product of the elastic modulus of the FRP shell in the hoop direction, E subscript FRP minus Hoop, times the wall thickness of the FRP tube, T subscript FRP, plus the quotient of 1 minus upsilon subscript C divided by the elastic modulus of concrete core, E subscript C.

Equation 2. The axial stress of the confined concrete core, F subscript CC, equals the quotient of the peak confined strength of the concrete, F prime subscript CC, times X times R, all divided by the total of R minus 1 plus X to the R power.

Equation 3. The term X equals epsilon subscript CC divided by the strain at peak strength F prime CC, which is epsilon prime CC.

Equation 4. The term F prime subscript CC equals F prime subscript C times the following: the product of 2.254 times the square root of the sum of 1 plus the quotient of 7.94 times sigma subscript R divided by F prime subscript C, all minus 2 times sigma subscript R divided by F prime subscript C, all minus 1.254.

Equation 5. The sample's relative density, D subscript R, equals 100 percent times the following quotient: the sum of the quotient of 1 divided by the sand's minimum dry unit weight, gamma subscript min minus the quotient of 1 divided by the dry unit weight of the sample, gamma subscript dry, all divided by the sum of the quotient of 1 divided by gamma subscript min minus the quotient of 1 divided by gamma subscript max.

Equation 6. Phi equals the sum of phi subscript O minus the product of delta phi times the log subscript 10 of the quotient of sigma prime subscript n divided by P subscript A.

Equation 7. The moisture content of the FRP specimen at submergence time T, which is M parentheses T, equals the moisture content at saturation, which is M subscript infinity, times the sum of 1 minus the summation from N equals 0 to infinity of the product of two terms. The first term is the quotient of 8 divided by the product of the sum, squared, of 2N plus 1, times pi squared. The second term is E to the following power: negative D, times the sum, squared, of 2N plus 1, times pi squared, times T divided by H squared.

Equation 8. The moisture content of the FRP specimen at submergence time T, which is M parentheses T, equals the moisture content at saturation, which is M subscript infinity, times the sum of 1 minus two products. The first product is the quotient of beta divided by alpha plus beta times E to the power of the product of negative alpha times T. The second product is the product of the quotient of beta divided by the sum of alpha plus beta times the summation from N equals 0 to infinity of the product of two terms. The first term is the quotient of 8 divided by the product of the sum, squared, of 2N plus 1, times pi squared. The second term is E to the following power: negative D, times the sum, squared, of 2N plus 1, times pi squared, times T divided by H squared.

Equation 9. The measured force on the Statnamic load cell, which is F subscript statnamic, equals the inertial resistance from effective mass of the foundation, which is F subscript inertia, plus the effective viscous damping resistance, which is F subscript damping, plus the effective soil resistance, which is F subscript static.

Equation 10. The mass moment of inertia about the base, which is I subscript Y, equals mass, M, times the following sum: the quotient of radius, R, squared, divided by 4, plus the quotient of the height, H, squared, divided by 3.

Equation 11. I subscript Y, equals mass, M, times the following sum: the quotient of the square pile of width, B, squared, divided by 12, plus the quotient of the height, H, squared, divided by 3.

Equation 12. F subscript inertia times Z equals the product of I subscript Y times the quotient of X with two dots above it divided by Z.

Equation 13. F subscript inertia equals the product of I subscript Y times the quotient of X with two dots above it divided by Z squared, all of which in turn equals the product of M subscript E times X with two dots above it.

Equation 14. The damping constant, D, equals a constant, C, divided by C subscript C, all of which in turn equals C divided by 2 times the square root of the product of the static stiffness, K, times M subscript E.

Equation 15. F subscript damping equals C times X with a dot over it, all of which in turn equals D times X with a dot over it times 2 times the square root of the product of K times M subscript E.

Equation 16. The estimated ultimate axial capacity, Q subscript T, equals the ultimate shaft capacity resulting from the surrounding soil in side shear, Q subscript S, plus the total ultimate tip load at the base or tip of the pile, Q subscript B, minus the weight of the pile, W subscript P.

Equation 17. The shaft side shear, which is Q subscript S, equals the integral with respect to depth, Z, taken over the range from 0 to the embedment depth of the pile, L, of the product of shear stress between the pile and the soil at a depth Z, which is F subscript S parentheses Z, times the pile perimeter at depth Z, which is P parentheses Z.

Equation 18. Q subscript b equals the tip or end area of the pile, A subscript B, times the total ultimate end-bearing or tip stress, Q subscript B.

Equation 19. The shaft shear stress, F subscript S, equals the product of the coefficient of lateral earth pressure, which is K subscript small delta, small delta being the interface friction angle between pile and soil, times the correction factor for K subscript small delta when small delta does not equal phi, which is C subscript F, times the free-field effective overburden pressure at depth Z, which is sigma prime subscript V parentheses Z, times the sine of small delta.

Equation 20. The ultimate end-bearing stress, Q subscript B, equals the product of the dimensionless factor dependent on pile depth-width ratio, which is alpha subscript T, times the bearing capacity factor, N prime subscript Q, times the free-field effective overburden pressure at the pile tip, sigma prime subscript V begin parenthesis Z equals L end parenthesis.

Equation 21. The term F subscript S parentheses Z equals the product of sigma prime subscript V parentheses Z times the earth pressure coefficient, K, times the tangent of the interface friction angle between pile and soil, which is small delta.

Equation 22. The term Q subscript B equals sigma prime subscript V begin parenthesis Z equals L end parenthesis times the bearing capacity factor, which is N subscript Q.

Equation 23. The peak local shaft friction, F subscript S, equals the product of the local radial effective stress at failure, which is sigma prime subscript RF, times the tangent of the failure or constant volume interface friction angle, which is small delta.

Equation 24. The local radial effective stress after installation, sigma prime subscript RC parentheses Z, equals the product of 0.029 times the CPT tip resistance, which is Q subscript C, times the quotient, raised to the 0.13 power, of sigma prime subscript V parentheses Z divided by P subscript A, times the quotient, raised to the negative 0.38 power, of H divided by R.

Equation 25. The increase in radial effective stresses, which is delta sigma prime subscript RD, equals the product of 2 times the operational shear modulus of sand, G, times the quotient of average radial displacement, which is small delta H, divided by R.

Equation 26. G equals CPT tip resistance, which is Q subscript C, divided by the following: the sum of A plus the product of B times eta minus the product of C times eta squared; and eta equals Q subscript C divided by the square root of the product of P subscript A times sigma prime subscript V.

Equation 27. Sigma prime subscript RF equals the sum of sigma prime subscript RC plus delta sigma prime subscript RD (for compression loading); sigma prime subscript RF equals the sum of the product of 0.8 times sigma prime subscript RC plus delta sigma prime RD (for tension loading).

Equation 28. The term Q subscript B equals the product of Q bar subscript C, which is the CPT tip resistance averaged over 1.5 pile diameters above and below the pile tip, times the following: 1 minus the product of 0.5 times the log subscript 10 of the quotient of the pile diameter, D, divided by the CPT diameter, D subscript CPT. The term Q subscript B also is greater than or equal to 0.13 times Q bar subscript C.

Equation 29. For Z less than or equal to Z subscript C, the unit side shear stress mobilized along a pile segment at movement Z, which is F subscript S, equals the product of the maximum unit side shear stress, which is F subscript S, max, times the quotient of the movement of pile segment, which is Z, divided by the movement required to mobilize F subscript S, max, which is Z subscript C. For Z greater than Z subscript C, F subscript S equals F subscript S, max.

Equation 30. For Z less than or equal to Z subscript C, the unit side shear stress mobilized along a pile segment at movement Z, which is F subscript S, equals the product of the maximum unit side shear stress, which is F subscript S, max, times the difference between two terms. The first term is the product of 2 times the square root of the quotient of the movement of pile segment, which is Z, divided by the movement required to mobilize F subscript S, max, which is Z subscript C. The second term is the quotient of Z divided by Z subscript C. For Z greater than Z subscript C, F subscript S equals F subscript S, max.

Equation 31. For Z less than or equal to Z subscript C, the pile tip resistance mobilized at displacement Z, which is Q, equals the product of the quotient, raised to the one-third power, of Z divided by Z subscript C, times the maximum tip resistance, which is Q subscript max. For Z greater than Z subscript C, Q equals Q subscript max.

Equation 32. The left side of this equation is the sum of four multicomponent terms. The first term is the product of the sum of tau plus the product of the derivative of tau with respect to R times DR, times the sum of R plus DR, times D theta, times DY. The second term is the negative of the product of tau, times R, times D theta, times DY. The third term is product of the sum of sigma subscript Y plus the product of the derivative of sigma subscript Y with respect to Y times DY, times the sum of R plus the quotient of DR divided by 2, times D theta, times DR. Final term is the negative of the product of sigma subscript Y, times the sum of R plus the quotient of DR divided by 2, times D theta, times DR. The right side of the equation is 0.

Equation 33. The derivative of tau times R with respect to R plus the product of R times the derivative of sigma subscript Y with respect to Y equals 0.

Equation 34. The derivative of tau times R with respect to R is approximately equal to 0.

Equation 35. The integral, from R subscript O to R, of the derivative in equation 34, that derivative being the derivative of tau times R with respect to R, equals 0, which leads to the following equation: the sum of the product of tau parentheses R, which is shear stress acting at a radial distance, T, from the centerline of the pile, times R,DS minus the product of tau parentheses R subscript O times R subscript O equals 0. Tau parentheses R equals the quotient of the product of tau parentheses R subscript O times R subscript O divided by R, all of which in turn equals the quotient of the product of tau subscript O times R subscript O divided by R.

Equation 36. The shear strain, gamma, equals the sum of the derivative of U with respect to Z plus the derivative of W with respect to R, all of which in turn is approximately equal to DW divided by DR.

Equation 37. Gamma parentheses R equals DW divided by DR, which in turn equals tau parentheses R divided by G parentheses R.

Equation 38. Gamma parentheses R equals DW divided by DR, which in turn equals the quotient of the product of tau subscript O times R subscript O divided by the product of R times G parentheses R. Thus DW equals the product of the quotient of the product of tau subscript O times R subscript O divided by G parentheses R times DR divided by R.

Equation 39. The term W subscript S equals the product of tau subscript O, times R subscript O, times the integral from R subscript O to R subscript M of DR divided by G times R, all of which in turn equals the product of the quotient of the product of tau subscript O times R subscript O divided by G, times the natural logarithm of the quotient of R subscript M divided by R subscript O.

Equation 40. The term R subscript M equals the product of 2.5, times the pile embedment depth, which is l, times the factor of vertical homogeneity of soil stiffness, which is rho, times the sum of l minus Poisson's ratio of the soil, which is nu.

Equation 41. The term K subscript S equals G divided by the product of R subscript O times the natural logarithm of the quotient of R subscript M divided by R subscript O.

Equation 42. Z subscript base equals the quotient of the product of the pile tip load, Q subscript B, times the sum of 1 minus nu squared, divided by the product of 2 times the elastic modulus of the soil beneath the pile tip, which is E subscript S, times the pile radius, which is R subscript O.

Equation 43. Z subscript base equals the quotient of the product of the pile tip load, which is Q subscript B, times the sum of 1 minus nu, divided by the product of 4 times G times the pile radius, R subscript O.

Equation 44. K subscript base equals the quotient of the product of 4 times G times R subscript O, divided by the sum of 1 minus nu.

Equation 45. G subscript O equals rho times the square of the shear wave speed, V subscript S.

Equation 46. The quotient of G subscript O divided by Q subscript C, the whole quotient with subscript Average, equals 1615 times the quotient, raised to the negative 0.764 power, of Q subscript C divided by the square root of sigma prime subscript VO.

Equation 47. The quotient of G subscript O divided by Q subscript C, the whole quotient with subscript Average, equals 1634 times the quotient, raised to the negative 0.75 power, of Q subscript C divided by the square root of sigma prime subscript VO.

Equation 48. G equals the quotient of CPT tip resistance, which is Q subscript C, divided by the following: A plus the product of B times nu, minus the product of C times eta squared; eta equals Q subscript C divided by the square root of the product of P subscript A times sigma prime subscript V.

Equation 49. Tau, the shear stress value corresponding to strain gamma equals shear strain, gamma, divided by the sum of two quotients: 1 divided by the initial tangent shear modulus, G subscript O, plus gamma divided by the shear stress value that the hyperbola tends asymptotically, tau subscript ult.

Equation 50. The maximum shear stress, tau subscript max, equals the failure ratio, R subscript F, times tau subscript ult.

Equation 51. G subscript sec equals G subscript O times the sum of 1 minus the quotient of the product of tau times R subscript F divided by tau subscript max.

Equation 52. G subscript sec equals G subscript O times the sum of 1 minus the quotient of the product of the quotient of the product of tau subscript O times R subscript O divided by R times R subscript F divided by tau subscript max.

Equation 53. The term W subscript S equals the product of the quotient of the product of tau subscript O times R subscript O divided by G subscript O times the natural logarithm of the quotient of a multicomponent term. The numerator of the term is the sum of the quotient of R subscript M divided by R subscript O minus the quotient of the product of tau subscript O times R subscript F divided by tau subscript max. The denominator of the term is 1 minus the quotient of the product of tau subscript O times R subscript F divided by tau subscript max.

Equation 54. The secant shear modulus, G subscript sec, divided by the initial shear modulus, G subscript O, equals 1 minus the product of F, an empirical curve fitting parameter, times the quotient, raised to the empirical curve fitting parameter G, of the current shear stress, tau, divided by the maximum shear stress, tau subscript max.

Equation 55. The term W subscript S equals the product of the quotient of the product of tau subscript O times R subscript O divided by the product of G subscript O times G, times the natural logarithm of a multicomponent term. The numerator of the term is the sum of the quotient, raised to the G power, of R subscript M divided by R subscript O minus the product of F times the quotient, raised to the G power, of tau subscript O divided by tau subscript max. The denominator of the term is 1 minus the product of F times the quotient, raised to the power of G, of tau subscript O divided by tau subscript max.

Equation 56. K superscript base subscript secant equals the product of K superscript base subscript O times the following: 1 minus the product of F times the quotient, raised to the G power, of Q subscript B divided by Q subscript B-max.

Equation 57. Z subscript base equals Q subscript B divided by the product of K superscript base subscript O times the following: 1 minus the product of F times the quotient, raised to the G power, of Q subscript B divided by Q subscript B-max.

Equation 58. Sigma M equals M plus DM minus M minus the product of V times DX plus the product of Q times DY minus the product of P times DX times the quotient of DX divided by 2, all of which in turn equals 0.

Equation 59. The second derivative of M with respect to X plus the product of Q times the second derivative of Y with respect to X minus the derivative of V with respect to X equals 0.

Equation 60. M equals the integral with respect to A, from A, of the product of sigma parentheses Z times Z.

Equation 61. The term U parentheses X,Y, which is the displacement in the X direction across the pile cross section, equals phi times X, which is the distance to the neutral plane, which in turn equals the product of the derivative of Y with respect to X, times Z. The term epsilon parentheses Z, which is strains in the X direction across the pile cross section, equals the derivative U with respect X, which in turn equals the product of the second derivative of Y with respect to X, times Z, all of which in turn equals kappa times Z. Sigma parentheses Z equals the product of E subscript P times epsilon parentheses Z, which in turn equals the product of E subscript P times kappa times Z.

Equation 62. M equals the integral with respect to A, from A, of the product of, begin parenthesis, E subscript P times kappa times Z, end parenthesis, times Z.

Equation 63. M equals the product of E subscript P times kappa times the integral with respect to A, from A, of Z squared, all of which in turn equals E subscript P times I subscript P times kappa, all of which in turn equals E subscript P times I subscript P times the second derivative of Y with respect to X.

Equation 64. The sum of the product of E subscript P times I subscript P times the fourth derivative of Y with respect to X, minus the derivative of V with respect to X, equals 0.

Equation 65. Sigma F subscript H equals the sum of the product of P parentheses X times DX minus DV, which in turn equals 0. Simplifying, the derivative of V with respect to X equals P parentheses X.

Equation 66. The sum of the product of E subscript P times I subscript P times the fourth derivative of Y with respect to X, minus P parentheses X, equals 0.

Equation 67. Delta equals the product of the quotient of the product of P times L raised to the fourth power divided by the product of 8 times EI, times the following: the sum of 1 plus the product of the quotient of E divided by the product of 2 times G, times the quotient, squared, of D divided by L.

Equation 68. The term Y subscript B equals B divided by 60. The term P subscript B equals the product of the quotient of B subscript S divided by A bar subscript S, times P subscript ult.

Equation 69. E subscript PY-max equals the product of K, which is a constant giving the variation of E subscript PY-max with depth, times X, which is the depth below ground surface.

Equation 70. The sum of the second derivative of the moisture concentration, C, with respect to R plus the product of the quotient of 1 divided by R times the derivative of C with respect to R equals the product of the quotient of 1 divided by the diffusivity for radial diffusion, which is D subscript R, times the derivative of C with respect to T.

Equation 71. C parentheses R, 0 equals 0.

Equation 72. C parentheses R subscript I, T equals C subscript I. C parentheses R subscript O, T equals C subscript O.

Equation 73. C parentheses R, T equals the sum of C parentheses R, T, plus C tilde parentheses R.

Equation 74. The sum of the quotient of D squared C tilde divided by DR squared plus the product of the quotient of 1 divided by R times the quotient of DC tilde divided by DR equals 0.

Equation 75. C tilde parentheses R subscript I equals C subscript I. C tilde parentheses R subscript O equals C subscript O.

Equation 76. C tilde parentheses R equals the product of the quotient of C subscript O minus C subscript I divided by the natural logarithm of the quotient of R subscript O divided by R subscript I, times the natural logarithm of the quotient of R divided by R subscript I, the whole product being added to C subscript I.

Equation 77. The term C parentheses R subscript I, T equals 0. The term C parentheses R subscript O, T equals 0.

Equation 78. The term C parentheses R, 0 equals the negative of C tilde parentheses R, which in turn equals the negative of the product of the quotient of C subscript O minus C subscript I divided by the natural logarithm of the quotient of R subscript O divided by R subscript I, times the natural logarithm of the quotient of R divided by R subscript I, the whole product being added to C subscript I before the negative is applied.

Equation 79. C parentheses R, T equals C tilde parentheses R plus the product of two multicomponent terms. The first term is the summation from M equals 1 to infinity of the product of E raised to a power times K subscript 0 parentheses beta subscript M, T. The power to which E is raised is the negative of the product of three subterms, the first of which is D subscript R, the second of which is the square of beta subscript M, and the third of which is T. The second multicomponent term is the integral with respect to R, and from R subscript I to R subscript O, of the product of R, times K subscript 0 parentheses beta subscript M, R, times C tilde parentheses R.

Equation 80. K subscript 0 parentheses beta subscript M, R equals the product of three multicomponent terms. the first term is pi divided by the square root of 2. The second term is a quotient. The numerator is the product of beta subscript M times J subscript 0, parentheses beta subscript M times R subscript O, times Y subscript 0, parentheses beta subscript M times R subscript O. The denominator is the square root of 1 minus the quotient of the square of J subscript 0, parentheses beta subscript M times R subscript O, divided by the square of J subscript 0, parentheses beta subscript M times R subscript I. The third term is the difference between two quotients. The first quotient is J subscript 0, parentheses beta subscript M times R, divided by J subscript 0 parentheses beta subscript M times R subscript O. From this quotient is subtracted the quotient of Y subscript 0, parentheses beta subscript M times R, divided by Y subscript 0, parentheses beta subscript M times R subscript O.

Equation 81. The quotient of J subscript 0 parentheses beta times R subscript I divided by J subscript 0 parentheses beta times R subscript O, minus the quotient of Y subscript 0 parentheses beta times R subscript I divided by Y subscript 0 parentheses beta times R subscript O equals 0.

Equation 82. C parentheses R subscript I, T equals 0. C parentheses R subscript O, T equals M subscript infinity.

Equation 83. C parentheses R subscript I, T equals M subscript infinity. C parentheses R subscript O, T equals M subscript infinity.