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1 INTRODUCTION

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The abutment areas around the existing Sack Dam have been periodically eroded during flood events, and various types of riprap and concrete debris have been used to backfill and protect the eroded areas. A scour hole exists immediately downstream of the existing Sack Dam near the left abutment.

The Poso Canal is owned by Central California Irrigation District (CCID), and crosses over the top of the Arroyo Canal through a concrete flume/bridge structure. The Arroyo Canal headworks structure is located approximately 300 feet downstream (west) of the Poso Canal crossing. The new fish screen will be located between the Poso Canal crossing and the Arroyo Canal headworks. The maximum canal water level considered in design of the Poso Canal is elevation 129.5 feet. Elevations reported within this report are in reference to the CCID vertical datum.

The CCID vertical datum was referenced to the NGS monument located on the Bridge/Poso Canal Flume at elevation 132.25 feet (CCID datum). The CCID vertical datum can be converted to NAVD88 using the following conversion equation:

CCID Datum (132.25 feet) � 2.2 feet = NAVD88.

The water level upstream of Sack Dam and within the approach to the Arroyo Canal headworks is approximately flat, but can vary depending on the flow being passed down the San Joaquin River and the Arroyo Canal. The top of the gates at the proposed dam will be set at a maximum height of elevation 122.0 feet (CCID vertical datum).

At any river flow greater than 45 cubic feet per second (cfs), selective pneumatic crest control gates on Sack Dam will start to lower to maintain a water surface elevation of 121.5 feet upstream of Sack Dam. When the San Joaquin River flow exceeds 1,550 cfs, the dam gates are completely down and the water level becomes uniform upstream and downstream of the dam. The maximum flood level considered in design of the Project is a river flow of 4,500 cfs, resulting in a water surface elevation of 125.6 feet at the proposed Sack Dam.

The critical design condition for the dam and canal structures is when there is a differential water height from upstream to downstream, creating lateral force on the structures from water pressure. The normal operating detained water level used to design Sack Dam is elevation 122.0 feet. For the trash rack structure, the maximum detained water level is elevation 126.0 feet (top of bulkhead) to account for flood conditions with the installed bulkheads. For the Fish Screen structure, the water will be approximately level across the screen except for minor head losses.

1.4 Borrow Sources for Fill Material

The right side levee of the Arroyo Canal has been built up with sediment dredged from the bottom of the canal over the years. The amount of soil available for fill and backfill of the new facilities varies from approximately 1 to 6 feet in thickness across the top of the levee. The soil material consists of fine sand, silt, and clay. The material is acceptable for construction, but may present challenges for achieving compaction. The soil contains clay balls with a high moisture content, and will require mixing or drying prior to compaction.

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SECTION 2

Field Exploration

2.1 Field Exploration

The field exploration was conducted during November, 2011, and consisted of excavating 8 test pits and drilling 10 soil borings at the locations shown on Figures 3 and 4. The locations shown are based on measurements from existing features using a fiberglass tape.

Four test pits were excavated within the east river overbank flow area where shown on Figure 3. The test pits were advanced to a maximum depth of 9 feet, but generally were terminated because of severe caving of the sidewalls. Four additional test pits were excavated within the right levee of the Arroyo Canal, where shown on Figure 4. The purpose of these test pits was to evaluate potential borrow materials for use as fill and backfill material. The east overbank flow area was being considered as a potential borrow source at the time of the field exploration, but has since been removed from consideration. The test pits were backfilled immediately after the materials and conditions encountered were recorded on test pit logs. The test pit logs are included in Appendix A.

The borings were drilled by Gregg Drilling of Martinez, California, using mud rotary methods to a maximum depth of 75 feet below existing ground surface (bgs). The borings within the San Joaquin River and Arroyo Canal were drilled using a CME 45 rig mounted on a barge, and the land borings were drilled using a Failing 1500 truck mounted rig. The floating barge was secured at the drill locations using steel spuds embedded into the riverbed.

In situ testing and soil sampling were performed in borings using the Standard Penetration Test (SPT). The SPT was performed in accordance with American Society of Testing and Materials (ASTM) D1586 at approximately 5 foot intervals. An automatic trip hammer was used. The SPT provides a disturbed sample of the soil and an empirical indication (N value) of the soil density or consistency. A modified California 2.5 inch, split spoon sampler was also used in an attempt to obtain intact samples for laboratory testing of a less disturbed specimen, preserved in brass liners placed within the sampler.

A Shelby tube sampler, which consists of a 3 inch inside diameter, thin wall steel tube, was also was used in fine grained soil materials that were encountered during the subsurface exploration. The steel walls are tapered and sharpened at the base to reduce disturbance to the soil during sampling. Relatively undisturbed samples were obtained by pushing the Shelby tubes for a distance of up to 2.5 feet into the soil, or until refusal, using the drill rig hydraulics.

All field exploration was directed by CH2M HILL geotechnical personnel. The soils encountered were visually classified in the field in general accordance with ASTM D2488 (visual manual procedure for description of soils).

The soil classification description and boring logs are presented in Appendix B. The boring logs summarize the methods used, soil sample intervals, sampling data, laboratory testing, soil types, and conditions encountered.

Where laboratory tests were available, the soil classification was also evaluated in accordance with the Unified Soil Classification System (ASTM D2487), and the soil classifications on the logs were updated if necessary.

The material types encountered are presented graphically on subsurface cross sections within Appendix C. These graphical logs are a simplification of the highly stratified soil material, and the boundaries between different soil types are for visual clarity only, and may not represent the stratifications that may be present between soil sample locations. The test pits are not shown on the graphical cross sections.

The elevation of the ground surface adjacent to each boring was approximated from topographic data from an aerial survey completed in 2011. The elevation of the ground surface or mudline under water was based on the measured water depth at the boring location subtracted from the water elevation that was observed at the existing Sack Dam, and assuming a level water surface from the dam throughout the Project area. The accuracy for all borings is approximately plus or minus 1 foot in elevation.

2 FIELD EXPLORATION

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2.2 Laboratory Testing

A laboratory test program was implemented to determine soil characteristics of the materials encountered within the soil borings and test pits. Testing included the following:

Atterberg limits (ASTM D4318) Sieve analysis (ASTM D422) Consolidation tests (ASTM D2435) Percent finer than No. 200 sieve (ASTM D1140) Moisture content and dry density (ASTM D2937) Consolidated drained direct shear test (ASTM D3080) Consolidated undrained triaxial shear tests with pore pressure measurements (ASTM D 4767) Moisture density relationship, or compaction curves (ASTM D1557) Unconfined compressive Strength (ASTM D2166) Hydraulic conductivity tests (ASTM D5084) Soil corrosivity suite of tests (minimum resistivity, pH, chloride, sulfate)

Laboratory test results are included in Appendix D.

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RDD\120730002 (CLR4904.DOCX) 3-1

SECTION 3

Geologic Setting

3.1 Regional Geology

The Project site is located near the center of the Great Central Valley Physiographic Province of California. The Great Central Valley is a northwest trending basin approximately 450 miles long by 50 miles wide. The Great Central Valley province is divided into the piedmont alluvial plain of the western San Joaquin Valley and the floodbasin of the San Joaquin River.

The floodbasin division underlies the Sack Dam Project area (Lettis, 1982). This division varies in width from 3 to 18 miles wide. The ground surface is gently sloped northwestward at an average gradient of approximately 3 to 5 feet per mile along the axis of the basin and is infilled with floodplain and river channel deposition. In the past under natural conditions, early summer snow melt from the Sierra Nevada to the east caused the San Joaquin River to overtop its banks and flood the basin. These floods are now contained within various impoundments in the foothills of the Sierra Nevada and flow is maintained within the river channels and canals by artificial levees.

The basin is bordered by Mesozoic platonic, volcanic, and metamorphic rocks of the Sierra Nevada mountains on the east and by Mesozoic and Cenozoic metamorphic and sedimentary rocks of the Coast Ranges on the west.

3.2 Local Geology

The Project area is underlain by Holocene deposits of the Dos Palos Alluvium (Lettis, 1982). Beneath Sack Dam and within the San Joaquin River channel, the alluvium is distinguished as generally coarse grained, arkosic channel deposits of Sierra Nevada provenance. Beneath the fish screen and canals, the alluvium is distinguished as fine to coarse grained, arkosic, terrace levee, point bar, and abandoned channel deposits of Sierra Nevada provenance.

The area to the east of the San Joaquin River channel is underlain by the Modesto Formation of Pleistocene geologic age (Lettis, 1982), which consists generally of fine grained, arkosic lower fan deposits of Sierra Nevada provenance.

3.3 Seismicity and Seismic Design Criteria

The Project is subject to ground shaking from earthquakes originating on various California faults and random earthquake sources. Review of the USGS fault and fold database for the Project vicinity (USGS, 2006) indicates that no faults are known to pass through the Project site. A number of fault systems are present at the perimeter of the Central Valley basin and in the adjacent ranges, generally running parallel in a northwesterly direction. The nearest active fault, defined as a fault with evidence of surface displacements within Holocene time, or within about the last 11,000 years, is the Ortigalita Fault, located approximately 19 miles west of the Project site. Active faults within approximately 60 miles of Sack Dam are summarized in Table 1. Characteristic seismic events originating on these active faults are unlikely to produce a peak ground acceleration (PGA) exceeding 0.16g at the Project site, based on a deterministic analysis of wave attenuation over the distance to the Project area, using the Boore Atkinson next generation attenuation (NGA) relationship (Boore et al., 2007). The probabilistic based PGA is slightly higher than the deterministic value as discussed below.

3 GEOLOGIC SETTING

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TABLE 1

Summary of Nearby Active Faults Geotechnical Exploration Report Arroyo Canal Fish Screen and Sack Dam Fish Passage Project

Fault Name Distance from Project (miles and direction)

Characteristic Magnitudea

Slip Rate (millimeters/year)a

Great Valley, Segment 9 19 west 6.6 1.5

Ortigalita 27 west 6.9 1

Foothills Fault System 32 east 6.5 0.01

Quien Sabe 42 west 6.4 1

Calaveras 45 west 6.4 15

San Andreas (creeping section) 47 west 6.5 34

Sargent 53 west 6.8 3

Zayante Vergeles 55 west 6.8 0.1

San Andreas (1906 Rupture segment) 57 west 7.9 24 a Characteristic magnitudes and slip rates are based on values either used by USGS to develop 2002 National Seismic Hazard maps or summarized in USGS, 1996.

A total of 45 earthquakes were identified in earthquake records from the last 200 years that had magnitudes greater than 5.0 (M5.0) within a 60 mile radius around the site. The largest earthquake to affect the Project area was the Coalinga Earthquake that occurred on May 2, 1983, with a magnitude of 6.4. The Coalinga Earthquake originated approximately 52 miles south of Sack Dam. The nearest earthquake event was a magnitude 5.0 event that occurred 31 miles southwest of the Project site in 1926 near Panoche Valley.

Potential strong ground motion at the Project site was estimated using probabilistic data from the national seismic hazard maps (USGS, 2008). The maps indicate that the PGA at the Project site with a 10 percent probability of exceedance in 50 years is 0.18g. This probability corresponds to a return period of about 475 years.

The mapped value is for a soft rock site condition (Site Class B), which needs to be adjusted for the soil site condition encountered at the Project site (a stiff soil site or Site Class D). A site amplification factor of 1.19 was used to result in a design PGA of 0.21g for the Project site.

The USGS deaggregation analyses showed that the dominant earthquake source contributing to the seismic hazard at the site is a magnitude 6.4 event at a distance of approximately 20 miles from Sack Dam. The source for the dominant earthquake was within the area of the Great Valley Fault.

For design of structures subjected to seismic loads, a soil site class of �D� may be used in accordance with the 2009 International Building Code (IBC). The site class was determined based on the standard penetration test (SPT) field N values for the materials encountered within the borings during the field exploration. The mapped spectral accelerations (Site Class B) for short periods and at 1 second period ground motion are 0.78g and 0.29g, respectively. These values should be corrected for site class �D� soil at the Project site to obtain ground surface design values for seismic loads in accordance with the IBC.

The bridges over the Poso Canal will be designed for seismic loads in accordance with Caltrans Seismic Design Criteria (Version 1.6, 2010). In general, the Design Spectrum is defined as the greater of a probabilistic spectrum based on a 5 percent in 50 years probability of exceedance (or 975 year return period), a deterministic spectrum based on the nearby faults, or a minimum spectrum considering a magnitude 6.5 earthquake on a strike slip fault located 12 kilometers from the Project. The deterministically controlling fault is the Great Valley Fault with a maximum earthquake magnitude of 7.0. The probabilistic spectrum exceeded both the minimum spectrum and the deterministic spectrum, with a peak ground acceleration of 0.27g and a maximum spectral acceleration of 0.58g at a period between 0.2 to 0.3 seconds. The site soil profile is Type D based on measured SPT values during site exploration, with an approximate shear wave velocity of 250 meters per second. The bridge structure is short

3 GEOLOGIC SETTING

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and will have a period less than 0.5 seconds; therefore, the standard 5 percent damped Acceleration Response Spectrum curve may be used without modification.

3.4 Subsidence

Subsidence has been ongoing within the San Joaquin valley due to groundwater withdrawal beginning in the mid 1920s, and increasing notably around 1945 (Galloway et al., 1999). In the period from 1926 to 1970, subsidence of approximately 2 feet in the vicinity of Sack Dam was estimated based on surveys by the U.S. Geological Survey (Ireland, 1983). The average rate of subsidence was between ½ to ¾ inch per year during this period.

Some areas of the San Joaquin Valley have experienced much greater rates of subsidence, such as 10 miles southwest of the town of Mendota, where close to 30 feet of subsidence was measured from 1926 to 1981.

Importation of surface water to the San Joaquin Valley began in the 1950s via the Friant Kern Canal and the Delta Mendota Canal. As a result, groundwater levels have recovered within many areas of the San Joaquin Valley, and subsidence has either slowed considerably or become negligible. The storage capacity of the aquifer has decreased due to consolidation, and the groundwater level quickly drops when pumping is resumed during drought years. During periods of drought in 1976 to 1977 and 1987 to 1991, subsidence was observed to increase when groundwater pumping was resumed temporarily.

More recently, in 2004, continuous operating reference stations were installed to monitor subsidence at the Mendota Dam, approximately 15 miles southeast of Sack Dam, and at check structure No. 18 on the Delta Mendota Canal, approximately 10 miles southwest of Sack Dam. Records provided by CCID indicate that these two areas have experienced subsidence of up to an average of 0.5 inch per year. The estimated subsidence from 1926 to 1970 at Mendota Dam was approximately 6 feet, or 1 to 2 inches per year, and at check structure No. 18 was approximately 12 feet, or approximately 3 to 4 inches per year (Ireland, 1983). The recent rate of subsidence appears to have slowed to less than 25 percent of the 1926 to 1970 rates.

For farms on the east side of the San Joaquin River near Sack Dam, surface water delivery is not available to the majority of lands adjacent to, or near, the river. Wells have recently been installed for irrigation, and a significant increase in subsidence in the vicinity of Sack Dam (more pronounced east of the San Joaquin River) has been measured from 2008 to the present time. A network of survey monuments is currently being monitored by the Bureau of Reclamation to record ongoing subsidence as part of the San Joaquin River Restoration Project. The subsidence rate at Sack Dam has been increasing since 2005, and appears to be subsiding at a rate of approximately 0.5 foot per year during the period from 2010 to 2011. The subsidence rate around the perimeter of the HMRD water delivery service area appears to be subsiding at a rate of 0.2 to 1.6 inches per year. If differential subsidence continues at the current rates, it is likely that water delivery will be affected significantly.

The study of subsidence in the vicinity of Sack Dam is ongoing, and conclusions and a plan of action will be developed as more data become available.

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SECTION 4

Subsurface Conditions

4.1 Soil Material

The Project area is underlain by alluvial deposits consisting of mixtures of fine to medium sand, silt, and clay soil materials. The soil is stratified in layers varying in thickness from a couple of inches to greater than 10 feet. Gravel was not encountered within the soil borings. Some hardened clay pieces up to ¾ inches in dimension were present and appeared as gravel on the soil gradation laboratory tests. One cobble up to 6 inches in dimension was observed at the test pit locations on the right side levee of the Arroyo Canal, deposited within the artificial fill material.

The sand material is variable in density, but is generally medium dense to dense except within the upper 10 feet of the Arroyo Canal. The silt and clay layers are generally stiff and over consolidated to some degree. There are localized zones of loose sand or soft clay, but these zones are not continuous between soil boring locations.

The majority of fines are of low plasticity, with a plasticity index less than 12. High plasticity soil, with a liquid limit greater than 50 or a plasticity index greater than 25, was not encountered.

Levees have been constructed along the Arroyo Canal, the Poso Canal, and the San Joaquin River. The levees within the areas explored consist of fine sand, silt, and clay soil, likely borrowed from the canal excavations based on the similarity to the native soil materials. The thickness or depth of artificial fill material on the levees can be approximated by comparing the elevation of the natural ground surface outside of the levees, and varies up to a maximum of approximately 10 feet. The clay material from within the test pit excavations was in the form of clay balls or �chunks� of clay.

4.2 Groundwater

Groundwater was encountered within the soil borings at depths near the water level within the San Joaquin River and Arroyo Canal. The water level within the San Joaquin River and Arroyo Canal varied during the field investigation because the water was being drained from the Mendota Pool, located approximately 23 miles upstream of the Project, during the week of drilling. The water level at Sack Dam was between elevation 119 and 120 feet during the field investigation.

A stabilized groundwater level within the soil borings was not possible to measure because the borings were full of drilling mud continuously, and the holes were grouted immediately after reaching final depth. Free water was visible in some of the soil samples, indicating saturated conditions, which provided an apparent groundwater level between the sampling intervals. The moisture condition is described on the soil boring logs (Appendix B) as �wet� where free water was observed in a soil sample. The degree of saturation was also calculated where moisture and density tests were determined, assuming a specific gravity of 2.7 for the soil, to estimate the phreatic water level.

The apparent water levels at the soil boring locations are shown on the subsurface cross sections included in Appendix C. The apparent groundwater level at the boring locations was within approximately 2 feet above and 5 feet below the water surface elevation in the river and Arroyo Canal. It is likely that saturated conditions above the water level in the river and canal were elevated by capillary action of the fine grained soil materials. In general, groundwater levels are expected to approximately match adjacent water bodies or irrigated fields.

There is a potential for groundwater levels to be elevated adjacent to the Poso Canal due to water infiltration from the unlined portions of the canal. The maximum water level within the Poso Canal is approximate elevation

129.5 feet.

4.3 Expansive Soil

High plasticity clay soil, or fat clay, was not encountered within the soil borings. Out of 46 Atterberg limit test results, only 2 tests had a plasticity index greater than 20. The risk of adverse consequences to the Project from

4 SUBSURFACE CONDITIONS

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expansive soil is low, based on typical ranges in Liquid Limit and Plasticity Index that exhibit expansive behavior (Nelson et al., 1992). No mitigation measures for expansive soil are required.

4.4 Soil Corrosivity

Laboratory testing was performed to determine the pH, resistivity, and chloride and sulfate contents of the native soil. One test was performed on shallow soil from 3 feet beneath the San Joaquin River bottom, and another from beneath the levee on the right side of the Arroyo Canal at a depth of 19 feet. The test results are included in Appendix D. Minimum resistivity of the two samples was low, chloride and sulfate content is significantly low relative to levels of concern, and pH is slightly alkaline. Based on the test results, the native soil should be considered moderately corrosive to buried ferrous metals due to low resistivity. The native soils are not expected to be corrosive toward buried concrete and cement mortar surfaces.

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SECTION 5

Analyses

5.1 Liquefaction Potential and Associated Hazards

Liquefaction is described as the sudden loss of soil shear strength due to a rapid increase of soil pore water pressures from cyclic loading during a seismic event. During seismic shaking, pore water pressure can build up until shear strength is significantly reduced. The potential adverse effects of liquefaction include ground settlement, ground cracking and expulsion of water and sand, the partial loss of bearing and confining forces used to support loads, increased lateral soil pressure, buoyancy, and lateral spreading.

Liquefaction occurs in soils located below groundwater level (saturated). Loose granular soils (sand, silty sand, sandy silt, and some gravels, usually associated with young geologic age) are particularly susceptible to liquefaction. Some clay and silt soils of low plasticity are also susceptible to liquefaction.

The alluvial soils encountered during the field exploration were generally dense or cohesive and not subject to liquefaction from the design ground acceleration (475 year return period), except for the upper 9 to 10 feet of soil material within the Arroyo Canal, where loose sand was encountered. Potential liquefaction of these loose sand deposits is discussed below.

The liquefaction potential was evaluated quantitatively based on SPT sampler penetration resistance (N values) observed in the drilled borings. Penetration resistance indicates the existing relative density or consistency of the underlying soil deposits, which correlates to liquefaction potential. The measured N values were corrected as described below to determine normalized blowcounts for use in the liquefaction analyses.

The liquefaction analyses were performed in accordance with the procedures that were developed by consensus of the participants of National Center for Earthquake Engineering Research workshops (Youd et al., 2001). The potential for liquefaction is estimated by calculating the estimated cyclic stress ratio induced by the design ground motion and compared with the capacity of the soil to resist liquefaction, expressed in terms of the cyclic resistance ratio. A maximum ground acceleration of 0.21g and an earthquake magnitude of 6.5 were used in the analyses, corresponding to the 10 percent probability of exceedance in 50 year ground motion and a characteristic magnitude earthquake event on the nearest faults, respectively. The risk of liquefaction is considered significant where the ratio of cyclic resistance ratio to cyclic stress ratio, or factor of safety, is less than 1.0.

The calculations for liquefaction potential are included in Appendix E. The factor of safety for liquefaction was calculated to be less than 1.0 for four soil samples out of all the soil sampling performed during the field investigation, three of which were within the upper 9 to 10 feet of soil within the Arroyo Canal between the Poso Canal crossing and the Arroyo Canal headworks. Liquefaction of the upper soil within the Arroyo Canal would lead to loss of bearing support and result in total or differential settlement for the fish screen and fish passage structures. The potential settlement of the fish screen in the event of the design earthquake was 5 inches, based on the methods of analysis presented by Tokimatsu and Seed (1987). To limit damage to the proposed fish screen structure, the liquefiable soil materials should be improved in place, replaced with properly compacted soil material, or the foundation should be deepened to be supported on non liquefiable soil materials below 10 feet depth using piles.

One sample from SB 3 on the north side levee of the Arroyo Canal, at a depth of 25 feet bgs, had a high risk of liquefaction for the design earthquake event. Liquefaction in limited pockets of loose sand material is not likely to cause instability to the proposed Project facilities; however, damage could occur to concrete slabs or asphalt concrete paving from resulting sand boils or localized ground depressions. The hard surfacing and any embankment displacement could require repair in the event that liquefaction were to occur.

Lateral spreading can occur when liquefied zones move downslope or areas overlying liquefied zones move downslope. Liquefiable soils were limited to the bottom of the Arroyo Canal and in a localized zone at 25 feet depth within soil boring SB 3, and are not likely to be continuously connected within the highly variable

5 ANALYSES

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subsurface materials. The risk level of damage to the proposed fish screen structure within the Arroyo Canal due to large scale lateral spreading is considered low.

5.2 Consolidation Settlement

A total of seven consolidation tests were performed for soil specimens collected during the field investigation.

Based on the laboratory test results, the preconsolidation pressure exceeded the existing overburden pressure in all of the test specimens.

The bearing pressures beneath the new Sack Dam, trash rack, and canal structures are relatively low. The bearing pressures are estimated to be between 400 to 900 pounds per square foot, well below the allowable bearing capacity of the soil. The estimated settlement beneath the center of these structures is approximately 1 inch due to consolidation settlement.

The largest potential for settlement is for the fill material on either side of the canal walls at the fish screen, and at the approaches to the trash rack structure. The estimated settlement of the fill material proposed within the Arroyo Canal between the canal structures and the existing levees due to consolidation of the underlying soil material is approximately 2.5 inches.

The time for consolidation depends on the thickness of the various clay layers. The soil is highly stratified, and it would be reasonable to estimate a maximum clay layer thickness of 10 feet, bounded by drainage layers at some point within that distance. Assuming drainage both up and down from a 10 foot thick clay layer, the time to achieve 90 percent primary consolidation settlement was estimated to range from 1 to 14 weeks for the various soils that were tested. For the consolidation settlement of the fill material on either side of the canal walls at the fish screen, half of the estimated total consolidation is anticipated to occur with 3 weeks after backfill is completed.

The risk of adverse affects on the proposed Project facilities from consolidation settlement is low, except for the control building and miscellaneous structures founded on the fill material. Some delay is likely between completion of backfill and construction of the control building. A delay of 6 weeks after completion of backfill is estimated to be sufficient to limit post construction settlement to less than approximately 1 inch, which should be considered in the structural design of the control building.

5.3 Stability and Seepage

5.3.1 Stability Analyses of River and Canal Structures

The Sack Dam and trash rack structures were analyzed for stability against sliding, overturning, and piping underneath the structures. The forces that are anticipated to act on the structures include the water pressure from the river on the dam and trash rack gates, uplift from internal water pressure beneath the structures, earthquake forces, and seepage forces. The upward seepage force at the downstream side of the structures, caused by a differential water head across the structure, can cause heaving of the soil if the seepage gradient is too large. The stability criteria used for analyzing the Sack Dam and Arroyo Canal structures are summarized in Table 2.

TABLE 2

Stability Criteria for River and Canal Structures Geotechnical Exploration Report Arroyo Canal Fish Screen and Sack Dam Fish Passage Project

Condition Sliding

Factor of Safety Overturning Factor of Safety

Minimum Base Area in Compression

(percent) Heave at Toe or

Piping Factor of Safety

Normal Operation 2.0 2.5 100 6

Maximum Design Flood 1.5 2.0 100 5

Earthquake 1.1 1.5 50 Not Analyzed

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The stability results are affected by the depth of the sheet pile cutoffs both upstream and downstream of the structures, because of their effect on the uplift pressure. The uplift pressure reduces the effective stress between the concrete floor slab and the subgrade soil, resulting in lower frictional resistance. A range in sheet pile cutoff depths was analyzed to evaluate the effect on the uplift pressure. The uplift pressure was determined from flow nets created for various sheet pile cutoff depths using the steady state finite element seepage model within the computer program SLIDE.

The frictional resistance along the base of the dam and trash rack is equal to the internal friction angle of the soil (33 degrees) times a friction reduction factor to account for the smoothness and hardness of the concrete. For the silt and sand soil, a friction reduction factor of approximately two thirds was assumed for the analyses. A friction angle of 23 degrees between the concrete and the alluvial soil material was used, or a friction factor of 0.4.

5.3.2 Loading Conditions and Resisting Forces for Stability Analyses The destabilizing loads on the Sack Dam and trash rack structures include horizontal water pressure, uplift pressure beneath the structures, and seismic forces from the design earthquake.

The proposed dam will have a potential to retain water to a maximum height of 5 feet above the top of the concrete floor slab, or elevation 122 feet. The normal operating water level with the gates up would be elevation

121.5 feet. The maximum water level considered for design is elevation 122 feet.

During a flood condition, the dam gates are lowered and the water level is relatively uniform across the dam from upstream to downstream, so static horizontal water pressure on the gates is approximately equalized and cancels out. The force of the river flow needs to be considered, however, during a flood condition because the water is moving rapidly downstream rather than being detained. For flood conditions, a maximum flow velocity of 5 feet per second was assumed. This would induce a dynamic horizontal force of approximately 250 pounds per foot of width on the dam structure, based on a water pressure equal to the water density times the flow velocity squared. The dynamic horizontal force is less than the maximum static horizontal water pressure force.

The trash rack design includes bulkhead gates with a potential to retain water upstream of the structure. The maximum retained water height is up to 12 feet above the floor slab, or elevation 126 feet, which would be the maximum detained water level during a flood condition before the water spills over the bulkheads. The retained water height under normal water level with the bulkhead gates in is elevation 122 feet, or 8 feet above the floor slab.

The fish screen does not have a mechanism to retain water.

Because of seepage losses around the sheet piles, the uplift pressure beneath the dam and trash rack is less than the retained water head, and was estimated from evaluation of a flow net. The uplift on the Sack Dam during a flood condition is negligible because of the increased height of water over the entire floor slab. The flow nets for the design water levels at the dam and trash rack are included within Appendix E.

For analyses under seismic loading, the lateral force on a rigid non building structure is equal to 0.3 times the 2009 IBC seismic design spectral response acceleration at short periods times the weight of the structure, times an importance factor of 1.25. The result is a lateral force equal to 23 percent of the structure weight that was applied for the seismic loading conditions. For allowable stress design methods, 70 percent of the lateral force is assumed to act on the structures, together with the full lateral water load.

The destabilizing loads are resisted by the weight of water above the upstream portion of the structures, and the weight of the structure itself. The upstream (for overturning resistance) or downstream (for sliding resistance) sheet piles also provide resistance to loads. A horizontal resisting force of 3,500 pounds per foot of width was included in the stability analyses, which is the load estimated to cause approximately 0.25 inch of deflection at the top of the sheet piles. A vertical resisting force of 2,500 pounds per foot of width was included in the stability analyses, estimated as the shear resistance of four 5/8 inch diameter steel anchors connecting the concrete slab to each pair of sheet piles.

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5.3.3 Seepage and Piping Potential

Seepage beneath the river or canal structures can lead to movement of soil particles, causing piping, where the seepage velocity in the microscopic flow channels beneath the structures becomes high enough to overcome the friction between the soil particles, and very fine soil will migrate towards an exit point. This process can become progressively worse until an ultimate failure of the structure occurs.

Piping generally tends to begin along the contacts between rigid structures and their adjacent soil foundations.

Line of creep theories have been developed by Bligh (1912) and were later refined by Lane (1935) based on experience with river structures on alluvial foundations. Analyses have been performed for the proposed new dam and canal structures using Lane�s line of creep theory, in which the ratio of length of seepage path to the water head difference across the structure is evaluated. The creep length is defined as the shortest path that a particle of water could take in flowing under a structure along the contact with the soil. The horizontal seepage path lengths are weighted less than the vertical creep lengths by a factor of one third due to higher hydraulic conductivity along horizontal paths.

It has been shown that the ratio of water head to the weighted creep length that is necessary to limit piping depends on the type of soil. The alluvial foundation soil at the Project consists of mainly fine to medium sand and silt. According to Lane�s recommendations, a minimum weighted creep ratio between 7 (for fine sand) and 8 (for silt) is required to prevent piping. A minimum weighted creep ratio of 8 was used for the design criteria for all river and canal structures.

Steel sheet pile cutoff walls or other water impediment barriers will be required both upstream and downstream of the new dam and trash rack.

5.3.4 Summary of Results

The stability results for the design loading conditions at the proposed Sack Dam and trash rack structure are summarized in Table 3. Stability calculations are included in Appendix E.

TABLE 3

Summary of Stability Analyses for Sack Dam and Trash Rack Structure Geotechnical Exploration Report Arroyo Canal Fish Screen and Sack Dam Fish Passage Project

Condition

Upstream Cutoff Deptha

Downstream Cutoff Depth

Sliding Factor of Safety

Overturning Factor of Safety

Percent of Base in

Compressionb

(percent)

Heave Factor of Safety

Lane�s Weighted Creep Ratio

Sack Dam

Sack Dam � Normal Retained Water Height, Static

35 20 3.7 2.7 100 19 24

Sack Dam � Normal Retained Water Height, with Earthquake

35 20 1.7 2.6 100 19 24

Arroyo Canal Trash Rack

Arroyo Canal Trash Rack � Normal Retained Water, Static

35 20 3.3 2.8 100 12 16

Arroyo Canal Trash Rack � Maximum Retained Water, Static

35 20 1.7 2.5 96 8 10

Arroyo Canal Trash Rack � Normal Retained Water, Earthquake

35 20 1.6 2.7 100 11 15 a Cutoff depths are measured from the top of the concrete floor slab.

b Bearing capacity was not critical for any loading conditions. The bearing pressures were less than 1,200 pounds per square foot for all of the loading cases.

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The earthquake loading is conservative because the weight of the concrete floor slab is included within the total weight of the structure when calculating the seismic force from self weight of the structure.

A sensitivity analysis was performed to evaluate the effect of leaky interlocks in the sheet piles on the stability of the structures. The result of leaky interlocks, modeled by reducing the hydraulic conductivity of the sheet piles to only 2 orders of magnitude less than the native soil, caused a reduction of approximately 10 percent in the factors of safety due to increased uplift forces. The stability of the structures does not appear to be jeopardized by leaking interlocks.

The proposed Sack Dam meets the design stability criteria for the sheet pile cutoff depths listed. The stability results for the trash rack retaining the maximum water level are slightly below the design criteria for percent of base in compression. These results are acceptable, however, because the trash rack will be supported on steel H piles, which will stabilize the structure against overturning, sliding, and differential settlement. The H piles were not included in the stability analyses.

5.4 Pile Capacity

5.4.1 Steel Pipe Piles for North and South Bridges

The pipe piles foundations for the bridges will be closed end, and backfilled with concrete, reinforced in the upper portion for connection to the bridge structure. An inside flange, conical pointed tip protector is recommended to aid in driving the pile through the stiff clay layers that were encountered at some boring locations. The piles are 16 inch diameter steel pipe piles.

The allowable vertical pile capacity was evaluated in accordance with the 2012 AASHTO LRFD Bridge Design Specifications. The factored load for the strength limit state controlled the required penetration depth of the pipe piles. The factored load for strength limit state was 44 kips per pile for the north bridge, and 81 kips for the south bridge, assuming a Caltrans permitted vehicle load of 54 kip maximum axle weight or a standard HS20 truck loading with a 32 kip maximum axle weight, whichever results in greater factored loads. The ultimate pile capacity for vertical loads was evaluated using the computer program DRIVEN Version 1.2, created by Blue Six Software of Logan Utah for the Federal Highway Administration. A layered soil profile of sand and clay material was used in the analyses to model the conditions encountered at soil borings SB 6 and SB 9. The upper 5 feet of soil was conservatively assumed to be very soft clay to account for the open canal and embankment slope adjacent to the piles.

The required penetration depth for the four north bridge piles at each abutment was 38 feet below the base of the pile cap. The required penetration depth for the four south bridge piles at each abutment was 53 feet below the base of the pile cap. The input parameters and resulting pile capacity are included in the program output files in Appendix E.

The lateral design earthquake loads on the bridges is 11.3 kips per pile at the north bridge and 21.1 kips per pile at the south bridge. The computer program LPile Plus version 5.0.47, created by Ensoft of Austin, Texas, was used to evaluate the performance and required depth of piles. A transformed (combined) pile section was used in the lateral pile analyses to account for the combined action of steel with concrete infill. According to the lateral load evaluation, the deflection under the earthquake loading did not significantly vary for pile depths greater than approximately 33 feet, and the vertical loads controlled the required depth of the piles. The maximum deflection, moment, and shear in the steel pipe piles were evaluated, and are presented in Appendix E along with the LPile Plus output results.

5.4.2 Steel Pipe Piles for Log Boom

Open ended pipe piles will support the log boom at the entrance to Arroyo Canal. The piles were analyzed for the combined load of the water flow force of 125 pounds per square foot and a floating debris impact load of 12 kips at the water surface level. The maximum river depth of 12 feet will result in the highest stress condition for the unbraced piles. Calculations of maximum displacement, moment, and shear within the piles are included in Appendix E. The purpose of these analyses was to estimate a recommended depth of embedment for the piles, 5-6 RDD\120730002 (CLR4904.DOCX)

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and confirm that the lateral moment and shear capacity are not exceeded for the proposed 18 inch diameter piles.

The computer program LPile Plus version 5.0.47 was used to evaluate the performance and required depth of piles. A minimum depth of 20 feet was required to limit excessive deflection of the pipe piles. A maximum deflection of approximately 4 inches is estimated for the design flood elevation of 126 feet. Including a corrosion allowance of 0.2 inch, the deflection increased to 11 inches at the top of the cantilevered, corroded piles. Final output computer files and summary graphs of deflection, moment, and shear in the piles are provided in Appendix E.

5.4.3 Steel H-Piles for Fish Screen Support

H piles will be used to provide support to the floor slab of the fish screen in the event of liquefaction occurring within the loose sand material underlying the fish screen area. The vertical pile capacity for HP 12x63 steel H piles was derived from a combination of side friction and end bearing. A pile penetration depth of 42 feet will support the estimated load of 80 kips per pile, with a factor of safety greater than 2. For the pile capacity analyses, the upper 10 feet of soil was considered to cause dragdown forces on the piles to account for the potential liquefaction. The analyses of pile capacity are included in Appendix E.

5.4.4 Steel H-Piles for Trash Rack Support

H piles will be used to provide lateral support to the trash rack structure from uneven loading resulting from water pressure upstream of the bulkhead gates during high river level at elevation 126 feet (NAVD88). The vertical pile capacity for steel H piles was evaluated assuming soil conditions encountered at soil boring SB 7.

The lateral load per pile, for a pile spacing of 10 feet on center, was 9 kips. The pile spacing is sufficient that it is not necessary to reduce the lateral capacity for group effects. The required H pile sections to resist the lateral loading was HP 12x63, with approximately 0.25 inch of displacement during the loading from the maximum design flood. The analyses of pile capacity are included in Appendix E.

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SECTION 6

Recommendations

6.1 Seepage Cutoff Depths

Based on the results of the stability and seepage analyses, the recommended minimum cutoff depths for steel sheet piles upstream and downstream of the proposed structures are summarized in Table 4. The fish screen does not have a significant differential water height across the screen; therefore, seepage beneath the fish screen structure is not a critical design issue. A sheet pile cutoff wall is recommended on the upstream side of the screen for erosion protection, but a cutoff wall on the downstream side of the screen is not necessary.

TABLE 4

Recommended Sheet Pile Cutoff Wall Tip Elevations Geotechnical Exploration Report Arroyo Canal Fish Screen and Sack Dam Fish Passage Project

Structure Upstream Tip Elevation

(feet) Downstream Tip Elevation

(feet)

Sack Dam 80 95

Trash Rack 80 95

Fish Screen Canal 94 None

6.2 Lateral Earth Pressures

Equivalent lateral pressures were estimated from the shear strength and density of the various soil materials that were tested. It is assumed that the backfill for various retaining walls will be borrowed from the right levee of the Arroyo Canal. The lateral pressures provided in Table 5 are applicable to either compacted fill or in situ native soil.

TABLE 5

Recommended Lateral Earth Pressures for Moist and Saturated Soil Geotechnical Exploration Report Arroyo Canal Fish Screen and Sack Dam Fish Passage Project

Equivalent Fluid Density (pcf)

Soil Density Active At rest Passive

Moist Native Soil, or Native used as Fill 40 60 390

Saturated Native Soil, or Native used as Fill 85 97 260

Earth Pressure Coefficients, Horizontal Backfill

Native Soil, or Native used as Fill Active: 0.31 At rest: 0.47 Passive: 3.2

Notes:

Typical Native Soil Properties: moist = 127 pcf; sat = 135 pcf; � = 32 degrees moist = moist unit weight sat = saturated unit weight � = effective internal friction angle pcf = pounds per cubic foot

1. Equivalent fluid pressures for saturated conditions include the hydrostatic pressure of the water in addition to the earth pressure for active, at rest, and passive conditions.

2. Compaction within 3 horizontal feet of walls should be performed with lightweight, hand operated equipment so that compaction induced lateral stresses are limited. If heavy or large equipment is used for compaction immediately adjacent to the structure, lateral stresses will be larger than those shown in this table.

6 RECOMMENDATIONS

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The lateral earth pressures provided in Table 5 do not include lateral forces induced from strong ground motion during earthquakes. For yielding walls (cantilever), a dynamic lateral force equal to 7H2 should be added to the static lateral active earth pressure, where H equals the height of the wall from the base of footing to the top of the wall. The dynamic lateral force should be applied at a height of 0.6H above the base of the wall footing. For non yielding walls (braced), the dynamic lateral force should be modeled as a uniform pressure equal to 7H.

Concrete footings and slabs cast against undisturbed native or structural fill soil may use a sliding friction factor of 0.40 in resisting lateral loads.