P907 TPP Pier Geotechnical Report.pdf

Geotechnical Engineering Services

P907 TPP Pier Naval Base Kitsap Bangor, Washington for The United States Navy and Notkin Mechanical Engineers

May 17, 2019

Geotechnical Engineering Services

Naval Base Kitsap Bangor, Washington for The United States Navy and Notkin Mechanical Engineers

May 17, 2019

17425 NE Union Hill Road, Suite 250 Redmond, Washington 98052 425.861.6000

May 17, 2019 | Page i File No. 0144-488-00

Table of Contents

1.0 INTRODUCTION

2.0 PROJECT DESCRIPTION

3.0 PREVIOUS STUDIES

4.0 FIELD EXPLORATIONS AND LABORATORY TESTING

4.1. Field Explorations

4.2. Laboratory Testing

5.0 SITE CONDITIONS

5.1. Geology

5.2. Seismicity

5.2.1. Earthquake Source Zones

5.2.2. Benioff Source Zone

5.2.3. CSZ Interplate Source Zone

5.2.4. Shallow Crustal Source Zone

5.3. Surface Conditions

5.4. Subsurface Conditions

5.4.1. Soil Conditions

5.4.2. Groundwater Conditions

6.0 CONCLUSIONS AND RECOMMENDATIONS

6.1. Summary of Geotechnical Considerations

6.2. Earthquake Engineering

6.2.1. Ground Rupture

6.2.2. Liquefaction

6.2.3. Recommended Response Spectrum

6.2.4. Lateral Spreading Induced Load on Piles

6.3. Deep Foundations

6.3.1. Axial Pile Capacity

6.3.2. TZPILE Parameters

6.3.3. LPILE Parameters

6.3.4. Construction Considerations

6.4. FLAC 2D Numerical Modeling Results

6.4.1. Seismically Induced Ground Displacements

6.4.2. Pile Responses Under Seismic Loading

6.5. Earthwork

6.5.1. Excavation Considerations

6.5.2. Stripping, Clearing and Grubbing

6.5.3. Subgrade Preparation

6.5.4. Erosion and Sedimentation Control

6.5.5. Structural Fill

6.5.6. Temporary Slopes

6.5.7. Permanent Slopes

6.6. Retaining Walls

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6.6.1. Design Recommendations

6.7. Drainage Considerations

6.7.1. Retaining Walls

6.7.2. Other Drainage Considerations

6.8. Pavement Recommendations

6.8.1. Subgrade Preparation

6.8.2. Asphalt Pavement

6.8.3. Portland Cement Concrete Pavement

6.9. Geotechnical Special Inspection

7.0 LIMITATIONS

8.0 REFERENCES

LIST OF FIGURES

Figure 1. Vicinity Map Figure 2. Site and Exploration Plan Figure 3. Cross Section A-A’ Figure 4. Cross Section B-B’ Figure 5. Cross Section C-C’ Figure 6. Recommended Site-Specific MCEr Response Spectrum (Nearshore) Figure 7. Recommended Site-Specific CLE (475-year) Response Spectrum Figure 8. Recommended Site-Specific OLE (72-year) Response Spectrum Figure 9. Lateral Spreading Load Distribution Diagram Figure 10. Maximum Depth of Lateral Spreading (feet) – OLE Event Figure 11. Maximum Depth of Lateral Spreading (feet) – CLE Event Figure 12. Maximum Depth of Lateral Spreading (feet) – MCE Event Figures 13 through 18. Axial Capacity Plots Figures 19 through 21. FLAC Model Geometry

APPENDICES

Appendix A. Previous Studies Appendix B. Field Explorations

Figure B-1 – Key to Exploration Logs Figures B-2 through B-11 – Log of Borings

Appendix C. Laboratory Testing Figures C-1 and C-10 – Sieve Analysis Results Figures C-11 and C-12 – Atterberg Limits Test Results

Appendix D. Results of Analytical Testing Appendix E. Site Specific Response Analysis

Figure E-1 – Cascadia Subduction Zone Intraslab Earthquake Basin Amplification Factors Figure E-2 – Cascadia Subduction Zone Interface Earthquake Basin Amplification Factors Figure E-3 – Source-Specific Basin Amplification Factors Figure E-4 – Rock Outcrop MCE UHS (Basin-adjusted) Figure E-5 – Rock Outcrop CLE (475-year) UHS (Basin-adjusted) Figure E-6 – Rock Outcrop OLE (72-year) UHS (Basin-adjusted) Figure E-7a – As-recorded Response Spectra MCE Event

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Figure E-7b – Spectrally Matched and Filtered (10Hz) Response Spectra, MCE Event Figures E-7c through E-7i – Acceleration, Velocity & Displacement Time Series Figure E-8a – As-recorded Response Spectra CLE Event Figure E-8b – Spectrally Matched and Filtered (10Hz) Response Spectra, CLE (475-year) Event Figures E-8c through E-8i – Acceleration, Velocity & Displacement Time Series Figure E-9a – As-recorded Response Spectra OLE Event Figure E-9b – Spectrally Matched and Filtered (10Hz) Response Spectra, OLE Event Figures E-9c through E-9i – Acceleration, Velocity & Displacement Time Series Figure E-10 – Nearshore Shear Wave Velocity Profile Figure E-11 – Recommended MCE Soil Amplification Factors Figure E-12 – Recommended CLE Soil Amplification Factors Figure E-13 – Recommended OLE Soil Amplification Factors Figure E-14 – Recommended Deterministic (MCER) Response Spectrum Figure E-15 – Recommended Site-Specific MCEr Response Spectrum (Nearshore) Figure E-16 – Recommended Site-Specfic CLE (475-year) Response Spectrum Figure E-17 – Recommended Site-Specfici OLE (72-year) Response Spectrum

Appendix F. Numerical Modeling FLAC Analyses Figures F-1 through F-5b – FLAC Model Geometry Figures F-6 and F-7 – Shear Wave Velocity Profile Figure F-8 – PM4Sand Calibration Procedure Figure Set F-9 – Lateral Load – Pile Head Displacements Figure F-10 – FLAC Pile Spring Calibration Load Settlement Curve Figures F-11 through F-13 – 2D Soil Only Model and 1D Site Response Spectra Comparison Figures F-14 through F-27– Horizontal Ground Displacement, Cross Section B-B’ and C-C’ Figure Sets F-28 through F-48. Pile Horizontal Displacement, Shear Force, Pile Moment, and

Axial Force Appendix G. Report Limitations and Guidelines for Use

May 17, 2019 | Page 1

1.0 INTRODUCTION

This report presents the results of GeoEngineers, Inc.’s (GeoEngineers’) subsurface explorations and geotechnical evaluation for the P907 Transit Protection Program (TPP) Pier and Maintenance Facility project at Naval Base Kitsap, Bangor, Washington. The project site is shown relative to surrounding physical features on the Vicinity Map (Figure 1) and the Site and Exploration Plan (Figure 2). This report provides geotechnical design recommendations for the 50 percent submittal for the new pier structure. The geotechnical engineering and design recommendations for the upland portions of the project including the new underground fuel facility and Vessel Maintenance Facility (VMF) are presented in a separate report.

The purposes of this report are to evaluate subsurface conditions at the proposed site and to provide geotechnical engineering and design recommendations for the proposed project components which include a new fixed pier with an access trestle.

GeoEngineers’ services were completed in general accordance with the Professional Services Subconsultant Agreement, Indefinite Quantity Contract, Project N44255-16-D-2006, between Notkin Mechanical Engineers and GeoEngineers, Inc. dated September 29, 2017 and Consultant Agreement-02 dated July 6, 2018. Our specific scope of services includes:

■ reviewing previous geotechnical explorations prepared for projects near the site;

■ completing offshore borings and supplemental offshore borings to characterize subsurface conditions for the site;

■ completing laboratory testing on selected soil samples obtained from the explorations;

■ completing numerical modeling of the soil conditions and pier piles;

■ developing geotechnical engineering design recommendations for the project; and

■ preparing this report.

2.0 PROJECT DESCRIPTION

GeoEngineers’ understanding of this project is based on the Facility Design Criteria (FDC) provided by the United States Navy (Navy) and discussions with the project team.

The TPP Pier will be located at Naval Base Kitsap (NBK) Bangor and will provide a berthing pier for two 250-foot blocking vessels. The pier will be located north of the existing K-B Docks pier and will be accessed via an approach trestle adjacent to the recently constructed Waterfront Operations Facility (Building 7136).

3.0 PREVIOUS STUDIES

As part of this evaluation, GeoEngineers’ reviewed available geotechnical reports completed as part of previous studies for projects located in the vicinity of the project area. The approximate locations of the previous explorations are shown on Figure 2. The reviewed geotechnical information includes:

■ The logs of borings GEI-1 and GEI-2 (2006) completed by GeoEngineers, Inc. for the geotechnical design of the P-990B Replace Buildings 7246/7247 Waterfront Operations Facility in 2013.

May 17, 2019 | Page 2

■ The logs of borings and S-1, B-24-3 (1973) completed by Shannon & Wilson, Inc. for the Trident Support Complex in 1973.

The boring logs from previous studies are presented in Appendix A.

4.0 FIELD EXPLORATIONS AND LABORATORY TESTING

4.1. Field Explorations

The subsurface soil and groundwater conditions at the site were evaluated by reviewing available geologic maps and drilling 13 borings (GEI-1 at the head of the trestle on land, GEI-2 through GEI-7 and GEI-11 and GEI-12 offshore for the pier). Borings GEI-2 through GEI-7 were completed on January 8 through January 14, 2018 using a barge (including boring GEI-5B which was drilled near GEI-5 after this boring was lost after the barge shifted). Boring GEI-1 was completed on February 5 and 7, 2018 using a track-mounted drill rig. Borings GEI-11 and GEI-12 were completed on August 21 to 24, 2018 using a barge. The borings were completed to depths of about 31 to 155 feet below mudline or ground surface. The borings were advanced with mud rotary drilling equipment.

The approximate locations of the borings completed for this project are presented on Figure 2. Details of the field exploration program and logs of the borings are presented in Appendix B. Details of the field screening methods and analytical testing is presented in Appendix D.

4.2. Laboratory Testing

Soil samples were obtained during the drilling program and taken to GeoEngineers’ laboratory for further evaluation. Selected samples were tested for the determination of moisture content, fines content (material passing the U.S. No. 200 sieve), grain size distribution (sieve analysis) tests, and Atterberg limits tests.

A description of the laboratory testing and the test results are presented in Appendix C.

5.0 SITE CONDITIONS

5.1. Geology

Published geologic information for the project vicinity includes the Washington State Department of Natural Resources (DNR) map titled “Geologic Map of the Seabeck and Poulsbo 7.5-minute quadrangles, Kitsap and Jefferson counties, Washington” (Polenz et al., 2013). The geologic map indicates that the surficial deposits in the vicinity of the pier consist of beach deposits. The units are described below in order of deposition, starting with the most recent.

■ Marine deposits consist of recent deposits of clay, silt, sand and gravel which are deposited in the near-shore environment. The marine deposits grade from loose to medium dense near the mudline to dense to very dense at depth.

■ Glacially consolidated soils consist of a dense to very dense, sand with variable silt and gravel content and very stiff to hard clay.

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5.2. Seismicity

5.2.1. Earthquake Source Zones

The Puget Sound area is located near the convergent continental boundary known as the Cascadia Subduction Zone (CSZ). The CSZ is the zone where the westward advancing North American Plate is overriding the subducting Juan de Fuca Plate. The interaction of these two plates results in two potential seismic source zones: (1) the Benioff source zone and (2) the CSZ interplate source zone. A third seismic source zone, referred to as the shallow crustal source zone, is associated with the north-south compression resulting from northerly movement of the Sierra Nevada block of the North American Plate.

5.2.2. Benioff Source Zone

Benioff source zone earthquakes are also referred to as intraplate, intraslab or deep subcrustal earthquakes. Benioff zone earthquakes occur within the subducting Juan de Fuca Plate between depths of 20 and 40 miles and typically have no large aftershocks. Extensive faulting results as the Juan de Fuca Plate is forced below the North American Plate and into the upper mantle.

The Olympia 1949 (M = 7.1), the Seattle 1965 (M = 6.5) and the Nisqually 2001 (M = 6.8) earthquakes are considered to be Benioff zone earthquakes. The Benioff zone is characterized as being capable of generating earthquakes up to magnitude 7.5. The recurrence interval for large earthquakes originating from the Benioff source zone is believed to be shorter than for the shallow crustal and CSZ source zones;

damaging Benioff zone earthquakes in Western Washington occur every 30 years or so. The deep focal depth of these earthquakes tends to dampen the shaking intensity when compared to shallow crustal earthquakes of similar magnitudes.

5.2.3. CSZ Interplate Source Zone

The CSZ is an approximately 650-mile-long thrust fault that extends along the Pacific Coast from mid-Vancouver Island to Northern California. CSZ interplate earthquakes result from rupture of all or a portion of the convergent boundary between the subducting Juan de Fuca Plate and the overriding North American Plate. The fault surfaces approximately 50 to 75 miles off the Washington coast. The width of the seismogenic portion of the CSZ interplate fault varies along its length. As the fault becomes deeper, materials being faulted become ductile, and the fault is unable to store mechanical stresses.

The CSZ is considered to be capable of generating earthquakes of magnitude 8 to 9. No earthquakes on the CSZ have been instrumentally recorded; however, through the geologic record and historical records of tsunamis in Japan, it is believed that the most recent CSZ event occurred in the year 1700 (Atwater 1996 and Satake et al. 1996). Recurrence intervals for CSZ interplate earthquakes are thought to be on the order of 400 to 600 years. Paleogeologic evidence suggests that five to seven interplate earthquakes may have been generated along the CSZ over the last 3,500 years at irregular intervals.

5.2.4. Shallow Crustal Source Zone

The shallow crustal source zone is used to characterize shallow crustal earthquake activity within the North American Plate. Shallow crustal earthquakes typically occur at depths ranging up to 12 miles. The shallow crustal source zone is characterized as being capable of generating earthquakes up to about magnitude 7.5. Large shallow crustal earthquakes are typically followed by a sequence of aftershocks. The largest known earthquakes associated with the shallow crustal source zone in Western Washington include

May 17, 2019 | Page 4 an event on the Seattle Fault about A.D. 900 and the 1872 North Cascades earthquake. The Seattle Fault event was believed to have been magnitude 7 or greater (Johnson et al. 1999), and the 1872 North Cascades earthquake is estimated to have been between magnitudes 6.8 and 7.4.

The site is located approximately 9 miles from the Seattle Fault zone (United States Geological Survey [USGS] 2010). The Seattle Fault Zone is a 2- to 4-mile-wide, west-trending zone of three or more south-dipping reverse faults (Johnson et al. 1999). The Seattle Fault ruptured about 1,100 years ago and caused broad uplift and subsidence on either side of the fault. The rate of recurrence on the Seattle Fault is thought to be on the order of thousands of years.

5.3. Surface Conditions

The proposed site for the TPP trestle/pier is situated west of Sealion Road about ½ mile south of the access trestle for Delta Pier. The proposed pier will be located near the tip of the K-B spit which projects into Hood Canal. The area at the head of the trestle has been graded relatively flat at about Elevation 17 to 19 feet (mean lower low water [MLLW]). A moderate to steep, 60- to 80-foot-high, slope separates the head of the trestle from Sealion Road which is situated above the shoreline at about Elevation 80 feet (northern end) to about Elevation 110 feet (southern end). An access road connects the pier site to Sealion Road which appears to have been cut into the existing slope with slopes as steep as 1.5H:1V (horizontal to vertical) above and below the road. An existing upper gravel parking lot is located on the west side of Sealion Road and slopes down towards the north from about Elevation 100 to 85 feet.

Numerous utilities are present near the head of the trestle including, but not limited to, storm/sewer drains, water, electrical and telecommunication lines.

5.4. Subsurface Conditions

5.4.1. Soil Conditions

The general site lithology consists of marine deposits overlying glacially consolidated soils. The marine deposits are divided into two categories; liquefiable and non-liquefiable. Boring GEI-1 was located at the head of the trestle and encountered marine deposits to a depth of approximately 43 feet below ground surface (bgs). Marine deposits in the offshore borings were thicker away from the shoreline and appear to be resultant of a longshore drift depositional environment due to currents within Hood Canal. The marine deposits were observed to consist of loose to very dense sand with variable silt, gravel and shell content.

Glacially consolidated soils were encountered below the marine deposits and consist of dense to very dense sand with variable gravel and silt content, dense to very dense gravel with variable sand and silt content, and very stiff to hard clay and silt. GeoEngineers’ interpretation of the subsurface conditions is presented in Cross Sections A-A’ through C-C’ (Figures 3 through 5).

5.4.2. Groundwater Conditions

Groundwater was observed at approximately 5 feet bgs (Elevation 9 feet) in boring GEI-1 at the head of the trestle. Groundwater is anticipated to vary in direct response to tidal changes in Hood Canal. No monitoring wells were installed. The offshore borings were completed using a truck-mounted drill rig located on a barge.

The barge was secured using anchors and/or spuds. The depth to each sample depth was determined using the relationship of the tide at the time of sampling, the length of the drill string, and periodic sounding.

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6.0 CONCLUSIONS AND RECOMMENDATIONS

6.1. Summary of Geotechnical Considerations

A summary of the primary geotechnical considerations is provided below. The summary is presented for introductory purposes only and should be used in conjunction with the complete recommendations presented in this report.

■ Site specific seismic hazard analyses were completed to develop site-specific design spectra for use in design of the trestle and pier structures. Recommended response spectra were developed at the risk-targeted maximum-considered earthquake (MCER), design earthquake (DE) (2/3 of MCER), contingency level earthquake (CLE), and operational level earthquake (OLE) levels, in general accordance with ASCE 61-14, Chapter 2 and ASCE 7-16, Chapter 21. Based on our review of the subsurface soil conditions at the site near the pier location, we conclude that after loose to medium dense marine deposits is susceptible to liquefaction under the design seismic events. For the MCE and CLE events, considerable thickness of the liquefiable soils will lose significant strength and stiffness, and the response of the pier structure will be dominated by the foundation bearing soils representative of the underlying dense to very dense soils. Therefore, the site-specific response spectrum for the design earthquake (MCE/DE) and the CLE were developed for the firm foundation bearing soil conditions, representative of site class C soil per ASCE 7. For the OLE event, the effects of the potentially liquefiable soils were considered to develop the design response spectra, representative of site Class E soil per ASCE 7.

■ The design team is planning to use 36-inch-diameter steel pipe piles to support the proposed trestle and pier structures. GeoEngineers has estimated axial capacities for the steel pipe piles and TZPILE parameters for use in the structural design. Lateral capacities of piles can be estimated using the LPILE soil parameters provided below.

■ Based on the results of our liquefaction-induced ground deformation analyses, we anticipate that seismic-induced lateral displacement will occur within the liquefiable portion of the marine deposits during the design seismic event. To account for the loads induced by seismic slope deformation, we recommend that the piles be designed for an additional kinematic soil pressure calculated using the earth pressure diagram presented herein.

■ Fast Lagrangian Analysis of Continua (FLAC) 2D numerical modeling was completed to estimate the thickness of laterally spreading soils that should be considered in the calculation of the kinematic load on the piles supporting the trestle and pier structures for the different earthquake levels considered for this project. In addition, the FLAC 2D model included the TPP pier and trestle structures to evaluate the pile responses under the DE events.

■ Based on our preliminary pile drivability analysis, we estimate that a Delmag D80 diesel impact hammer would be suitable to drive the proposed piles to the design tip elevation. Because the pile contractor has control of the pile/hammer configuration and the driving equipment, we recommend that the pile contractor be made responsible for selecting the appropriate pile-driving hammer and installing the piles to design tip elevation without damaging the piles.

These geotechnical issues and other considerations are discussed in greater detail, and conclusions and recommendations for the geotechnical aspects of the project are presented in the following sections.

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6.2. Earthquake Engineering

6.2.1. Ground Rupture

Because of the thickness of the glacially consolidated soils below the site and the estimated distance to the closest known fault is about 9 miles, it is our opinion that the potential for surface fault rupture is low.

6.2.2. Liquefaction

Liquefaction is a phenomenon where soils experience a rapid loss of internal strength as pore water pressures increase in response to strong ground shaking. The increased pore water pressure may temporarily meet or exceed soil overburden pressures to produce conditions that allow soil and water to flow, deform, or erupt from the ground surface. Ground settlement, lateral spreading and/or sand boils may result from soil liquefaction. Structures supported on or within liquefied soils may suffer foundation settlement or lateral movement that can be damaging.

The evaluation of liquefaction potential is a complex procedure and is dependent on numerous site parameters, including soil grain size, soil density, site geometry, static stresses, and the design ground acceleration. Typically, the liquefaction potential of a site is evaluated by comparing the cyclic shear stress ratio (the ratio of the cyclic shear stress to the initial effective overburden stress) induced by an earthquake to the cyclic shear stress ratio required to cause liquefaction. Estimation of the cyclic shear stress required to initiate liquefaction and the cyclic shear stress initiated by a DE was completed using empirical methods.

The cyclic shear stress ratio required to cause liquefaction at the site was estimated using empirical procedures based on correlations from the standard penetration tests (SPTs). Estimated ground settlement resulting from earthquake-induced liquefaction was analyzed using an empirical procedure that relates settlement to average SPT N-values. This analysis assumes a level ground surface.

In general, soils that are susceptible to liquefaction at this site include very loose to medium dense marine deposits that are below the groundwater table. Based on our analyses, the marine deposits encountered near the trestle abutment, trestle and pier structures are moderately to highly susceptible to liquefaction under the DE event. The medium dense to very dense marine deposits and very dense glacial soils have a low potential for liquefaction. Should the marine deposits liquefy and lose strength, both settlement and lateral deformation (Lateral Spreading) of the site towards the Hood Canal may occur.

We evaluated liquefaction potential of the site soils for three earthquake events: OLE, CLE, and maximum-considered earthquake (MCE) events per the American Society of Civil Engineers (ASCE) 61-14 Code. Our liquefaction analyses were completed using the boring information completed at the project site.

We evaluated liquefaction potential using the simplified method proposed by Youd et al (2001) and Idriss and Boulanger (2008). The seismic design parameters used in our liquefaction analyses are provided in Table 1 below.

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TABLE 1. SEISMIC DESIGN PARAMETERS FOR LIQUEFACTION ANALYSES

Design Earthquake Magnitude

Peak Ground Acceleration, PGAM (g)

OLE 6.9 0.20

CLE 7.6 0.41

MCE 7.9 0.55

Notes:

OLE - Return Period of 72-years CLE - Return Period of 475-years MCE - Return Period of 2,475-years

Based on our analyses, we estimate that the factor of safety is less than 1 during the design-level earthquakes for most of the deposits above a depth of about 20 feet for the trestle abutment, up to about 45 feet on the west end of the trestle structure, and between about 40 to 45 feet within the pier footprint.

Liquefaction-induced free-field ground settlement of the potentially liquefiable zones is estimated to be on the order of 3 to 31 inches for the design-level earthquakes. The magnitude of liquefaction-induced ground settlement will vary as a function of the characteristics of the earthquake (earthquake magnitude, location, duration and intensity) and the soil and groundwater conditions.

6.2.3. Recommended Response Spectrum

The recommended response spectra were developed at the MCER, DE (2/3 of MCER), CLE, and OLE earthquake levels, in general accordance with ASCE 61-14, Chapter 2 and ASCE 7-16, Chapter 21.

Based on our review of the subsurface soil conditions at the site near the pier location, we conclude that after the MCE/DE and the CLE events, the liquefiable soils will lose significant strength/stiffness and are expected to flow away. The response of the pier structure will be dominated by the foundation bearing soils representative of site class C soil per ASCE 7. Therefore, the site-specific response spectra for the design earthquake (MCE/DE) and CLE were developed for the firm foundation bearing soil conditions. For the OLE event, the effects of the potentially liquefiable soils were considered to develop the design response spectra, representative of site Class E soil per ASCE 7.

We completed site-specific seismic hazard analyses to develop the design spectra for the MCE, DE, CLE, and OLE levels for use in the structural design for this project as presented in Tables 2 through 5 and Figures 6 through 8. The recommended site-specific response spectra were developed at the ground surface including soil amplification effects. We understand that the recommended site-specific response spectra will be used for structural design based on the structural vibration periods and for the deformation analysis. The approach of developing the site-specific response spectra is described in detail in Appendix E, Site-Specific Response Analysis.

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TABLE 2. RECOMMENDED SITE-SPECIFIC TABLE 3. RECOMMENDED SITE-SPECIFIC

MCER RESPONSE SPECTRUM DE (2/3 MCER) RESPONSE SPECTRUM

Period (sec) 5% Damped Spectral Acceleration, Sa (g)

Period (sec)

Acceleration, Sa (g)

0.01 0.625 0.01 0.417

0.05 1.094 0.05 0.729

0.075 1.259 0.075 0.839

0.09 1.337 0.09 0.892

0.2 1.337 0.2 0.892

0.3 1.337 0.3 0.892

0.4 1.337 0.4 0.892

0.45 1.337 0.45 0.892

0.75 0.795 0.75 0.530

1 0.596 1 0.398

2 0.298 2 0.199

3 0.199 3 0.133

4 0.149 4 0.099

5 0.119 5 0.080

7.5 0.064 7.5 0.042

10 0.036 10 0.024

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TABLE 4. RECOMMENDED SITE- TABLE 5. RECOMMENDED SITE-SPECIFIC

SPECIFIC CLE RESPONSE SPECTRUM OLE RESPONSE SPECTRUM

Period (sec)

Acceleration, Sa (g)

Period (sec)

Acceleration, Sa (g)

0.01 0.277 0.01 0.203

0.05 0.521 0.05 0.414

0.075 0.591 0.075 0.476

0.1 0.611 0.1 0.496

0.2 0.744 0.2 0.606

0.3 0.628 0.3 0.622

0.4 0.531 0.4 0.383

0.5 0.454 0.5 0.238

0.75 0.333 0.75 0.141

1 0.233 1 0.088

2 0.082 2 0.024

3 0.050 3 0.013

4 0.031 4 0.008

5 0.024 5 0.006

7.5 0.012 7.5 0.003

10 0.008 10 0.002

6.2.4. Lateral Spreading Induced Load on Piles

Lateral spreading involves lateral displacements of large volumes of liquefied soil. Lateral spreading can occur on near-level ground as blocks of surface soils are displaced relative to adjacent blocks. Lateral spreading also occurs as blocks of surface soils are displaced toward a nearby slope or free-face by movement of the underlying liquefied soil. The Hood Canal shoreline represents a sloping condition. Lateral spreading of the trestle abutment area could occur during a design level earthquake should the marine deposits liquefy, resulting in the movement of soil towards the Hood Canal. In addition, lateral spreading could result in the movement of soil or sediment onto below-water piles.

Based on the results of our liquefaction-induced ground deformation analyses, we anticipate that seismic-induced lateral displacement will occur along the Hood Canal during the design seismic event.

To account for the loads induced by seismic slope deformation, we recommend that the piles be designed for an additional kinematic soil pressure calculated using the earth pressure diagram presented in Figure 9.

Resultant forces can be calculated from the earth pressure distribution. Figures 10 through 12 show the thickness of laterally spreading soils that need to be considered in the calculation of the kinematic load on the piles supporting the trestle and pier structures for the OLE, CLE and MCE events, respectively. These depths were determined based on the results of our soil liquefaction analysis and FLAC numerical modeling results (described in the FLAC section below). Note that the earth pressure diagram presented in Figure 9 is a generalized earth pressure developed for the entire site and we anticipate that it will be refined based on the shear and moment computed in our FLAC 2D models, as presented in the FLAC section below.

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6.3. Deep Foundations

The design team is planning to use 36-inch-diameter steel pipe piles to support the proposed trestle and pier structures. GeoEngineers has estimated axial capacities for the steel pipe piles, as described below.

The following sections also provide discussion of general construction considerations.

6.3.1. Axial Pile Capacity

Given the large diameter of the steel pipe piles, we estimated the axial pile capacities assuming unplugged conditions at the pile tip at the end of driving. Figures 13 through 15 show the static downward and uplift capacities for the trestle abutment, trestle, and pier structures, respectively. Figures 16 through 18 show the seismic downward and uplift capacities for the trestle abutment, trestle, and pier structures, respectively. The allowable capacities shown include a factor of safety of 2.0.

The capacities apply to single piles. If piles are spaced at least three pile diameters on center, as recommended, no reduction for group action is needed. The structural characteristics of pile materials and structural connections may impose limitations on pile capacities and should be evaluated by the structural engineer.

6.3.2. TZPILE Parameters

We understand that BergerABAM will incorporate the axial soil-pile interaction effects using the computer program TZPILE to design the planned trestle and pier piles. Our recommended soil parameters for the TZPILE analysis are presented in Tables 6 through 8 below, for the trestle abutment, trestle, and pier structures, respectively.

TABLE 6. TZ-PILE SOIL PARAMETERS – TRESTLE ABUTMENT

Soil Unit Soil Model

Reference Depth (ft)

Soil Parameters

Effective Unit

Weight (pcf)

Friction Angle (deg)

Undrained Shear

Strength (psi) E50

Ultimate Unit Side Friction

(psi) (top/bottom)

Ultimate Unit Tip Resistance

(psi) (top/bottom)

Liquefiable Marine Deposits

Driven Pile in Sand (Mosher)

0 to 10* 47.6 6 - - 0.034/0.312 0.075/0.68

Non- Liquefiable Marine Deposits

Driven Pile in Sand (Mosher)

10* to 30 52.6 35 - - 2.33/7.64 175/573

Glacially Consolidated Soils

Driven Pile in Clay (Coyle & Reese)

30 to 90 62.6 - 12/32 0.005 19.6/65 1518/5016

Notes:

*Depth= 10 feet represents MCE maximum lateral spreading depth provided in the Figure 12 pcf = pounds per cubic feet psi = pounds per square inch

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TABLE 7. TZ-PILE SOIL PARAMETERS – TRESTLE

Soil Unit Soil Model

Reference Depth (ft)

Soil Parameters

Effective Unit

Weight (pcf)

Friction Angle (deg)

Undrained Shear

Strength (psi) E50

Ultimate Unit Side Friction

(psi) (top/bottom)

Ultimate Unit Tip Resistance

(psi) (top/bottom)

Liquefiable Marine Deposits

Driven Pile in Sand (Mosher)

0 to 10* 47.6 6 - - 0.034/0.312 0.075/0.68

Non- Liquefiable Marine Deposits

Driven Pile in Sand (Mosher)

10* to 55 52.6 35 - - 2.33/14.64 175/1098

Glacially Consolidated Soils

Driven Pile in Clay (Coyle & Reese)

55 to 90 62.6 - 20/32 0.005 37/63 2859/4899

Notes:

*Depth= 10 feet represents MCE maximum lateral spreading depth provided in the Figure 12 pcf = pounds per cubic feet psi = pounds per square inch

TABLE 8. TZ-PILE SOIL PARAMETERS – PIER

Soil Unit Soil Model

Reference Depth (ft)

Soil Parameters

Effective Unit

Weight (pcf)

Friction Angle (deg)

Undrained Shear

Strength (psi) E50

Ultimate Unit Side Friction

(psi) (top/bottom)

Ultimate Unit Tip Resistance

(psi) (top/bottom)

Liquefiable Marine Deposits

Driven Pile in Sand (Mosher)

0 to 10* 47.6 6 - - 0.034/0.312 0.075/0.68

Non- Liquefiable Marine Deposits

Driven Pile in Sand (Mosher)

10* to 45 52.6 35 - - 5/23 175/888

Glacially Consolidated Soils

Driven Pile in Clay (Coyle & Reese)

45 to 90 62.6 - 16/32 0.005 30/64 2323/4946

Notes:

*Depth= 10 feet represents MCE maximum lateral spreading depth provided in the Figure 12 psi = pounds per square inch

May 17, 2019 | Page 12

6.3.3. LPILE Parameters

We understand that BergerABAM will incorporate the lateral soil-pile interaction effects using the computer program LPILE to design the planned trestle and pier piles. Our recommended soil parameters for the LPILE analysis are presented in Tables 9 through 11 below, for the trestle abutment, trestle, and pier structures, respectively.

TABLE 9. L-PILE SOIL PARAMETERS – TRESTLE ABUTMENT

Soil Unit Soil Model Reference Depth (ft)

Soil Parameters

Effective Unit Weight

(pcf)

Friction Angle (deg)

Cohesion (psf)

K (pci) E50 P-Multiplier

Liquefiable Marine Deposits

Sand (Reese) 0 to 10* 47.6 6

(residual) - 60 - 0.1

Non- Liquefiable Marine Deposits

Sand (Reese) 10* to 30 52.6 35 125 - -

Glacially Consolidated Soils

Silt (cemented c-phi)

30 and below 62.6 40 400 200 0.005 -

Notes:

*Depth= 10 feet represents MCE maximum lateral spreading depth provided in the Figure 12 psf = pounds per square foot pcf = pounds per cubic feet pci = pounds per cubic inch

TABLE 10. L-PILE SOIL PARAMETERS – TRESTLE

Soil Unit Soil Model Reference Depth (ft)

Soil Parameters

Effective Unit Weight

(pcf)

Friction Angle (deg)

Cohesion (psf)

K (pci) E50 P-Multiplier

Liquefiable Marine Deposits

Sand (Reese) 0 to 10* 47.6 6

(residual) - 60 - 0.1

Non- Liquefiable Marine Deposits

Sand (Reese) 10* to 55 52.6 35 125 - -

Glacially Consolidated Soils

Silt (cemented c-phi)

55 and below 62.6 40 400 200 0.005 -

Notes:

*Depth= 10 feet represents MCE maximum lateral spreading depth provided in the Figure 12 psf = pounds per square foot pci = pounds per cubic inch

May 17, 2019 | Page 13

TABLE 11. L-PILE SOIL PARAMETERS – PIER

Soil Unit Soil Model Reference Depth (ft)

Soil Parameters

Effective Unit Weight

(pcf)

Friction Angle (deg)

Cohesion (psf)

K (pci) E50 P-Multiplier

Liquefiable Marine Deposits

Sand (Reese) 0 to 10* 47.6 6

(residual) - 60 - 0.1

Non- Liquefiable Marine Deposits

Sand (Reese) 10* to 45 52.6 35 125 - -

Glacially Consolidated Soils

Silt (cemented c-phi)

45 and below 62.6 40 400 200 0.005 -

Notes:

*Depth= 10 feet represents MCE maximum lateral spreading depth provided in the Figure 12 psf = pounds per square foot pcf = pounds per cubic feet pci = pounds per cubic inch

Piles spaced closer than five pile diameters apart will experience group effects that will result in a lower lateral resistance for trailing rows of piles with respect to leading rows of piles for an equivalent deflection.

We recommend that the lateral load capacity for piles in a pile group spaced less than five pile diameters apart be reduced in accordance with the factors in Table 12 per American Association of State Highway and Transportation Officials (AASHTO) Load and Resistance Factor Design (LRFD) Bridge Design Specifications Section 10.7.2.4.

TABLE 12. PILE P-MULTIPLIERS, PM, FOR MULTIPLE ROWS

Pile Spacing1 (In Terms of Pile Diameter)

P-Multipliers, Pm2, 3

Row 1 Row 2 Row 3 and Higher

3D 0.80 0.40 0.30

5D 1.00 0.85 0.70

Notes:

1 The P-multipliers presented are a function of the center-to-center spacing of piles in the group in the direction of loading expressed in multiples of the pile diameter, D.

2 The values of Pm were developed for vertical piles only.

3 The P-multipliers are dependent on the pile spacing and the row number in the direction of the loading. To establish values of Pm for other pile spacing values, interpolation between values should be conducted.

May 17, 2019 | Page 14

To account for the variations in lateral pile resistance due to sloping ground, we recommend that P-multipliers be applied to the LPILE parameters in Tables 9 through 11 when evaluating lateral resistance of the trestle and pier piles. The illustration below provides the appropriate P-multiplier to be used for specific cases of directional lateral loading of piles along the slope. P-multipliers need only be applied for analyses of soil layers located within the dimension “X” shown below.

6.3.4. Construction Considerations

The piles for the proposed trestle and pier structures should be installed using an appropriately sized pile-driving hammer. The pile-driving hammer should be of sufficient size to drive the piling to the design tip elevation without damaging the pile. We anticipate that vibratory hammer will be used to drive the piles to practical refusal or 5 feet above the design tip elevation prior to driving the remainder of the piles to the design tip elevation. Based on our preliminary pile drivability analysis, we estimate that the minimum pile hammer that should be used to drive the proposed piles to the design tip elevation is Delmag D100 diesel impact hammer or equals. Because the pile contractor has control of the pile/hammer configuration and the driving equipment, we recommend that the pile contractor be made responsible for selecting the appropriate pile-driving hammer and installing the piles to design tip elevation without damaging the piles.

The depths and thicknesses of the interpreted soil units vary across the site. If pile resistance encountered during driving indicates that the soil conditions may differ significantly from those assumed for design, it may be necessary to install a longer pile or additional piles to achieve the recommended axial and lateral capacity used in the design. We therefore recommend that a monitoring program be implemented for the pile-driving operations.

GeoEngineers should be retained to observe the pile driving and to evaluate driving records to determine whether the soil conditions encountered during pile installation are consistent with those assumed for final design. If soil conditions are significantly different from those assumed, it will be appropriate for GeoEngineers to develop revised design criteria.

6.4. FLAC 2D Numerical Modeling Results

We completed FLAC 2D numerical modeling that incorporated the structural elements of the TPP pier site to evaluate the pile responses under the OLE, CLE and MCE events. We completed the numerical modeling for two representative cross sections (Cross Sections B-B’ and C-C’) to evaluate the magnitude of lateral slope and ground displacements at the P907 TPP Pier site under the DE events. Interpreted soil conditions

May 17, 2019 | Page 15 for Cross Sections B-B’ and C-C’ are shown in Figures 4 and 5, respectively. Three Soil-Structural Models (Model Trestle, Model Pier B-B’ and Model Pier C-C’) were developed to incorporate piles at different locations within the offshore structures based on the drawings provided by BergerABAM dated November 2, 2018.

Figures 19 through 21 present the three FLAC 2D models developed for this project. For each of the DE levels, we selected three earthquake ground motions for use as the input ground motions in our FLAC 2D model as presented in Table 13 below. Prior to propagating through the FLAC model, the three selected seed ground motions were first modified via spectral matching to match the target rock outcrop MCER response spectrum from 0.01 to 10 seconds using RSPMatch09 (Fouad, et al. 2012; Al Atik, et al. 2010).

TABLE 13. CHARACTERISTICS OF INPUT EARTHQUAKE TIME HISTORIES

Event Earthquake Source Mechanism Mw Station Component

OLE (72-year) Loma-Prieta, 1989 Crustal

(Reverse) 6.9 San Jose – Santa Teresa Hills 225

Tarapaca, 2005 Subduction-intraslab 7.8 Iqueque EW

CLE (475-year) Loma-Prieta, 1989 Crustal

(Reverse) 6.9 San Jose – Santa Teresa Hills 225

Maule, 2010 Subduction-interface 8.8 Concepcion San Pedro NS

DE (2,475-yr)

Loma-Prieta, 1989 Crustal (Reverse) 6.9 Gilroy Array #3 Fault Normal

Maule, 2010 Subduction-interface 8.8 Concepcion San Pedro NS

Tohoku, 2011 Subduction-interface 9.0 Onoda MYGH05 EW

The FLAC 2D modeling results include the seismically induced ground displacements of the submerged slopes and the pile responses (deflections, shear and moment) supporting the TPP piers and trestle. Details of our FLAC 2D analysis and results are presented in Appendix F, Numerical Modeling FLAC Analysis. The following presents a summary of the FLAC 2D modeling results.

6.4.1. Seismically Induced Ground Displacements

Our FLAC 2D numerical modeling results were used to estimate the thickness of laterally spreading soils that are recommended to be considered in the calculation of the kinematic load on the piles supporting the trestle and pier structures for the different earthquake levels considered for this project, as presented in Figures 10 through 12.

6.4.2. Pile Responses Under Seismic Loading

Pile deflected shapes, shear forces, moments and axial forces at the end of the earthquake were computed and are presented in Appendix F.

May 17, 2019 | Page 16

In our FLAC 2D dynamic analysis (Soil-Structural Models), we tracked the time histories of pile displacements, shear forces, moments and axial forces at different pile elevations. The maximum pile response during the OLE, CLE and MCE earthquake events are presented in Tables 14A through 14C. The maximum pile response at the end of the earthquake for the OLE, CLE and MCE events are presented in Tables 15A through 15C.

TABLE 14A. MAXIMUM PILE RESPONSES DURING EARTHQUAKE FOR OLE EVENT

Model Pile Displacement (in) Shear Force (kips) Moment (kips-ft) Axial Force (kips

Trestle

J (Plumb) 0.3 10 118 261

H-R (Batter) 0.3 6 102 573

H (Plumb) 0.3 5 93 421

H-L (Batter) 0.3 3 57 606

G-R (Batter) 0.3 5 88 522

G (Plumb) 0.3 4 75 381

G-L (Batter) 0.3 4 63 550

F-R (Batter) 0.3 6 99 494

F (Plumb) 0.3 5 86 396

F-L (Batter) 0.3 4 57 566

E-R (Batter) 0.3 8 130 311

E (Plumb) 0.3 6 108 205

E-L (Batter) 0.3 5 68 370

Pier B-B’

D-R (Batter) 0.4 5 39 288

D-L (Batter) 0.4 5 43 374

C (Plumb) 0.4 5 56 569

B (Plumb) 0.4 8 53 608

A (Plumb) 0.4 10 74 356

Pier C-C’

D-R (Batter) 0.3 4 68 334

D-L (Batter) 0.3 5 48 344

C (Plumb) 0.3 4 61 598

B (Plumb) 0.3 5 42 577

A (Plumb) 0.3 6 72 367 Notes:

In the pile column, ‘-R’ indicates pile batter direction up slope and ‘-L’ indicates pile batter direction down slope.

May 17, 2019 | Page 17

TABLE 14B. MAXIMUM PILE RESPONSES DURING EARTHQUAKE FOR CLE EVENT

Model Pile Displacement (in) Shear Force (kips) Moment (kips-ft) Axial Force (kips

Trestle

J (Plumb) 0.9 20 262 287

H-R (Batter) 1.0 21 299 948

H (Plumb) 0.9 21 306 492

H-L (Batter) 0.9 14 232 1109

G-R (Batter) 1.0 19 288 847

G (Plumb) 0.9 20 291 428

G-L (Batter) 0.9 17 260 1025

F-R (Batter) 1.0 21 320 830

F (Plumb) 0.9 25 322 435

F-L (Batter) 0.9 21 300 1026

E-R (Batter) 1.0 27 376 606

E (Plumb) 0.9 24 359 235

E-L (Batter) 0.9 26 352 751

Pier B-B’

D-R (Batter) 1.3 26 99 638

D-L (Batter) 1.3 24 204 886

C (Plumb) 1.3 26 201 600

B (Plumb) 1.3 29 130 653

A (Plumb) 1.3 30 190 382

Pier C-C’

D-R (Batter) 0.9 16 185 767

D-L (Batter) 0.9 13 134 839

C (Plumb) 0.9 16 197 647

B (Plumb) 0.9 18 105 626

A (Plumb) 0.9 22 202 394

In the pile column, ‘-R’ indicates pile batter direction up slope and ‘-L’ indicates pile batter direction down slope.

May 17, 2019 | Page 18

TABLE 14C. MAXIMUM PILE RESPONSES DURING EARTHQUAKE FOR MCE EVENT

Model Pile Displacement (in) Shear Force (kips) Moment (kips-ft) Axial Force (kips

Trestle

J (Plumb) 2.1 42 555 343

H-R (Batter) 3.0 126 603 1502

H (Plumb) 1.9 42 646 502

H-L (Batter) 1.9 37 498 1658

G-R (Batter) 2.9 80 534 1396

G (Plumb) 1.8 45 567 434

G-L (Batter) 1.9 39 500 1665

F-R (Batter) 2.8 49 541 1357

F (Plumb) 1.7 42 587 454

F-L (Batter) 1.9 39 538 1647

E-R (Batter) 2.5 47 597 1096

E (Plumb) 2.1 48 582 342

E-L (Batter) 1.9 40 478 1208

Pier B-B’

D-R (Batter) 4.0 65 208 1282

D-L (Batter) 3.4 33 286 1511

C (Plumb) 3.4 37 293 699

B (Plumb) 3.7 47 233 711

A (Plumb) 3.3 44 273 433

Pier C-C’

D-R (Batter) 2.2 29 409 1521

D-L (Batter) 2.2 26 312 1550

C (Plumb) 2.2 31 422 724

B (Plumb) 2.4 33 240 698

A (Plumb) 2.2 40 334 437

In the pile column, ‘-R’ indicates pile batter direction up slope and ‘-L’ indicates pile batter direction down slope.

May 17, 2019 | Page 19

TABLE 15A. MAXIMUM PILE RESPONSES AT THE END OF EARTHQUAKE FOR OLE EVENT

Model Pile Displacement (in) Shear Force (kips) Moment (kips-ft) Axial Force (kips

Trestle

J (Plumb) 0.1 3 22 248

H-R (Batter) 0.1 4 48 355

H (Plumb) 0.1 3 35 400

H-L (Batter) 0.1 1 12 347

G-R (Batter) 0.1 4 50 285

G (Plumb) 0.1 2 34 354

G-L (Batter) 0.1 1 11 355

F-R (Batter) 0.1 4 67 275

F (Plumb) 0.1 4 50 369

F-L (Batter) 0.1 1 26 384

E-R (Batter) 0.1 8 104 116

E (Plumb) 0.1 5 84 176

E-L (Batter) 0.1 4 60 203

Pier B-B’

D-R (Batter) 0.1 5 21 50

D-L (Batter) 0.1 2 20 187

C (Plumb) 0.1 5 35 533

B (Plumb) 0.1 7 33 582

A (Plumb) 0.1 10 69 328

Pier C-C’

D-R (Batter) 0.02 3 10 87

D-L (Batter) 0.03 1 7 132

C (Plumb) 0.02 1 6 575

B (Plumb) 0.02 3 13 554

A (Plumb) 0.03 5 45 346

In the pile column, ‘-R’ indicates pile batter direction up slope and ‘-L’ indicates pile batter direction down slope.

May 17, 2019 | Page 20

TABLE 15B. MAXIMUM PILE RESPONSES AT THE END OF EARTHQUAKE FOR CLE EVENT

Model Pile Displacement (in) Shear Force (kips) Moment (kips-ft) Axial Force (kips

Trestle

J (Plumb) 0.3 11 107 256 H-R (Batter) 0.4 18 198 233 H (Plumb) 0.3 18 191 412 H-L (Batter) 0.3 13 173 514 G-R (Batter) 0.4 18 240 172 G (Plumb) 0.3 18 216 368 G-L (Batter) 0.3 16 217 510 F-R (Batter) 0.4 20 277 192 F (Plumb) 0.3 23 245 373 F-L (Batter) 0.3 20 243 540 E-R (Batter) 0.3 26 324 31 E (Plumb) 0.2 23 303 169 E-L (Batter) 0.3 23 347 294

Pier B-B’

D-R (Batter) 0.5 23 56 53 D-L (Batter) 0.5 21 112 308 C (Plumb) 0.5 23 100 525 B (Plumb) 0.5 24 58 580 A (Plumb) 0.5 26 114 309

Pier C-C’

D-R (Batter) 0.1 11 33 68 D-L (Batter) 0.1 10 28 168 C (Plumb) 0.1 11 39 578 B (Plumb) 0.1 16 57 558 A (Plumb) 0.1 20 114 336

In the pile column, ‘-R’ indicates pile batter direction up slope and ‘-L’ indicates pile batter direction down slope.

May 17, 2019 | Page 21

TABLE 15C. MAXIMUM PILE RESPONSES AT THE END OF EARTHQUAKE FOR MCE EVENT

Model Pile Displacement (in) Shear Force (kips) Moment (kips-ft) Axial Force (kips

Trestle

J (Plumb) 0.6 16 116 259

H-R (Batter) 1.9 125 229 362

H (Plumb) 0.4 21 257 293

H-L (Batter) 0.4 25 246 541

G-R (Batter) 1.8 77 286 278

G (Plumb) 0.3 31 294 286

G-L (Batter) 0.4 29 289 619

F-R (Batter) 1.6 40 310 255

F (Plumb) 0.3 28 308 310

F-L (Batter) 0.4 28 331 659

E-R (Batter) 1.3 37 386 42

E (Plumb) 0.1 37 377 269

E-L (Batter) 0.4 29 396 309

Pier B-B’

D-R (Batter) 1.6 54 137 76

D-L (Batter) 1.3 23 97 313

C (Plumb) 1.2 23 94 614

B (Plumb) 1.5 29 68 585

A (Plumb) 1.2 28 161 317

Pier C-C’

D-R (Batter) 0.4 12 62 63

D-L (Batter) 0.4 11 35 168

C (Plumb) 0.5 12 48 585

B (Plumb) 0.8 19 92 565

A (Plumb) 0.5 25 107 335 Notes:

In the pile column, ‘-R’ indicates pile batter direction up slope and ‘-L’ indicates pile batter direction down slope.

6.5. Earthwork

Earthwork will be limited to the end of the trestle where it connects to the existing K-B Spit. The majority of the project will consist of pile installation offshore.

6.5.1. Excavation Considerations

Loose/soft to medium dense beach deposits were observed in the explorations within the anticipated excavation depths. We anticipate that these soils can be excavated with conventional excavation equipment, such as trackhoes or dozers.

6.5.2. Stripping, Clearing and Grubbing

We recommend that all new foundation, slab and pavement areas be stripped of existing foundations and pavements/structures. Any remaining below-grade elements from previous site development should also be removed. Abandoned, below-grade utilities should be removed; alternatively, below-grade utilities can be abandoned in place by completely filling the conduits with lean concrete.

May 17, 2019 | Page 22

Vegetation including the root mass and organic-rich topsoil should be stripped and removed from the sidewalk and pavement areas.

Depressions/excavations that result from removal of existing improvements that are present in new sidewalk or pavement areas should be filled (if located in areas where proposed grades are higher than the base of the depression/excavation) with properly compacted structural fill.

6.5.3. Subgrade Preparation

The exposed subgrade should be evaluated after site grading is complete. Proof-rolling with heavy, rubber-tired construction equipment should be used for this purpose during dry weather. Probing should be used to evaluate the subgrade during periods of wet weather. Soft areas noted during proof-rolling or probing should be excavated and replaced with compacted structural fill. Working pads consisting of nonwoven geotextile fabric and quarry spalls/crushed rock may be necessary to create a stable working platform.

6.5.4. Erosion and Sedimentation Control

Potential sources or causes of erosion and sedimentation depend upon construction methods, slope length and gradient, amount of soil exposed and/or disturbed, soil type, construction sequencing and weather.

Implementing an erosion and sedimentation control plan will reduce the project impact on erosion-prone areas. The plan should be designed in accordance with applicable city, county and/or Navy standards.