ML20041G415
| ML20041G415 | |
| Person / Time | |
|---|---|
| Site: | Wolf Creek  |
| Issue date: | 04/03/1981 |
| From: | SARGENT & LUNDY, INC. |
| To: | |
| Shared Package | |
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| References | |
| SL-3831, NUDOCS 8203220228 | |
| Download: ML20041G415 (150) | |
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'I NUCLEAR SAFETY-REL ATED I ENGINEERING DATA COMPILATION l FOR WATER-CONTROL STRUCTURES AT WOLF CREEK LAKE WOLF CREEK GENERATING STATION 1
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I REPORT PREPARED FOR I KANSAS GAS & ELECTRIC COMPANY I
l AND KANSAS CITY POWER & LIGHT COMPANY g
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nEPORT SL-3831 APRIL 3.1981 I ;= -
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April 3,1981 (
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s Mr. M. L. Johnson, Manager Nuclear Plant Engineering '
Kan:,as dos and Electric Company ,
~ - P.O, Box?08 201 North Market Street Wichita, Kansas 67201 '
Dear Mr. Johnson:
. s Enclosed for your use are 12 cocies of the following report:
Report SL-383L - _. i. ,
Engineerinir Data Compilation for Water-Con +rol Structures ( ,
Wolf Creek Genertting Station The purpose of this report is to comply'with the NRC Regulatory Guide 1.127 re-qutrement for compilation of all engin'ecring data related to water-control struc-
,tures which are specifically built for use iq conji'n:ti3n with a nuclear power plant anri whose failure could cause radiologhal consequences adversely affecting public
- 1 health and safety. This report presents the engineering data for the Wolf Creek ultimate heat sink dam, ultimate heat sink, main darn service spillway, and main dam auxiliary spillway.
Very truly yours,
. M. McLaughh,i Manager Structural Department JMM/RDN/rg Enclosures Copics:
D. Crawford (1/1)
P. J. Conroy (1/1) l l
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I NUCLEAR SAFETY-RELATED l ENGINEERING DATA COMPILATION FOR WATER-CONTROL STRUCTURES AT WOLF CREEK LAKE g
WOLF CREEK GENERATING STATION I
I I
lI REPORT PREPARED FOR KANSAS GAS & ELECTRIC COMPANY l AND KANSAS CITY POWER & LIGHT COMPANY I
I REPORT SL-3831 APRIL 3.1981
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I l l TABLE OF CO N T E N TS g
04-03-81 ruoc
1.0 INTRODUCTION
1 1.1 Purpose of the Report 1 1.2 Purpose of the Project 1 2.0 UHS DAM AND UllS 2 2.1 Ilydrologic and liydraulic Design 2 2.1.1 Design Criteria 2 2.1.2 Design and Analysis 2 2.1.2.1 Capacity of the UliS 2 2.1.2.2 Sedimentation of the UllS 3 2.1.3 Reference Drawings 3 Geotechnical Design 3 I 2.2 2.2.1 Design Criteria 3 2.2.1.1 Dam Geometry 3 2.2.1.1.1 UllS Dam 3 2.2.1.1.2 UllS Slopes 4 2.2.1.1.2.1 Slope Characteristics 4 2.2.1.1.2.2 Description of UHS Slopes 4 2.2.1.2 Soils 5 2.2.1.3 Filter Material 5 Gradation 6 I 2.2.1.3.1 2.2.1.3.2 Thickness 7 2.2.1.4 Riprap Material 7 I 2.2.1.4.1 2.2.1.4.2 Gradation Thickness 7
7 2.2.1.5 Compaction 8 2.2.1.6 Safety Factors - Stability Analysis 8 2.2.2 Geotechnical Data 8 2.2.2.1 Soil and Rock Investigation 8 ii SL-3831 I
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04-03-81 PAGE 2.2.2.1.1 Subsurface Exploration 8 2.2.2.1.2 Testing 9 2.2.2.2 Seismology 10 2.2.3 Design Analysis 10 2.2.3.1 Filter and Riprap Materials 10 2.2.3.1.1 General 10 2.2.3.1.2 Materials for UliS Dam 10 Materials for Slopes Adjacent to ESWS Pumphouse 12 I 2.2.3.1.3 2.2.3.2 Crest Width 12 2.2.3.3 Slope Stability Analysis 13 2.2.3.3.1 General 13 2.2.3.3.2 Ultimate lleat Sink Dam (UllSD) Stability 13 2.2.3.3.2.1 Shear Strength of Mcterials 13 2.2.3.3.2.2 Static Stability Analysis 14 2.2.3.3.2.3 Seismic Stability Analysis 15 2.2.3.3.2.3.1 Pseudostatic Analysis 15 Finite Element Analysis 16 I
2.2.3.3.2.3.2 2.2.3.3.2.3.2.1 Introduction 16 2.2.3.3.2.3.2.2 Design Earthquake and Loading Conditions 17 2.2.3.3.2.3.2.3 Procedure Used in Seismic Stability Evaluation 17 2.2.3.3.2.4 Stability Analysis Using Static Strength Following Cyclic Loading 21 J
2.2.3.3.3 Stability of ESWS Facility Slopes 21 2.2.3.3.3.1 Natural Slopes 21 2.2.3.3.3.2 Man-Made Slopes 22 2.2.3.3.3.2.1 Ultimate lient Sink Slopes 22 2.2.3.3.3.2.2 ESWS Pumphouse Intake Channel Slopes 23 2.2.3.3.3.2.3 Slopes Near the Discharge Structure 24 2.2.3.4 Settlement 24 2.2.3.5 Camber 25 2.2.3.6 Seepage Control 25
- 2. 2. 3.7 Solution and Weathering 26 iii SL-3831 I
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04-03-81 I PAGE 2.2.3.8 Liquefaction Potential 27 2.2.4 Instrumentation and Monitoring Program 27 2.2.4.1 Settlement Points on UllS Dam 27 2.2.4.2 Sedimentation Monitoring Points in UllS Pond and Intake Channel 28 2.2.4.3 Profile 28 2.2.4.4 Visual Observations 28 2.2.5 Reference Drawings 28 3.0 SERVICE SPILLWAY 29 3.1 Hydrologic and Hydraulic Design 29 3.1.1 Design Criteria 29 3.1.2 Design and Analysis 29 3.1.3 Reference Drawings 31 3.2 Structural Design 32 3.2.1 Design Criteria 32 3.2.2 Materials and Allowable Stresses 32 3.2.3 Design and Analysis 32 3.2.3.1 Weir 32 3.2.3.2 Wingwalls at Weir End 33 3.2.3.2.1 General 33 Sliding Check 33 lg 3.2.3.2.2 5 3.2.3.2.3 Overturning Check 34 l 3.2.3.2.4 Section Design for Shenrs and Moments 34 3.2.3.3. Wingwalls at Stilling Basin End 36 3.2.3.4 Channel 37 3.2.3.5 Bridge 39 3.2.4 Reference Drawings 40 4.0 AUXILIARY SPILLWAY 41 4.1 flydrologic and Ilydraulic Design 41 4.1.1 Design Criteria 41 4.1.2 Design and Analysis 41 iv SL-3831
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PAGE 4.1.3 Reference Drawings 42 4.2 Structural Design 42 4.2.1 Design Criteria 42 4.2.2 Reference Drawings 43
5.0 REFERENCES
44 APPEN DIX Appendix A - Description of Computer Programs Referenced in the Report l
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I LIST OF T A BLES 04-03-81 2-1 Required Safety Factors - Slope Stability Analysis, Category I Structures 2-2 Characteristics of Onsite Aggregate Sources 2-3.1 Soil Parameters for Stability Analysis - UHS Dam 2-3.2 Representative Soil Properties for Use in Design Analysis 2-3.3 Effective Stress Parameters - Modified Mohr Diagram 2-4 Results of Unconfined Compression Tests on Rock Core Samples 2-5 Results of Slope Stability Analyses for UHS Dam 2-6 Soil Parameters for Static Stress Analysis of Submerged UHS Dam 2-7 Initial Stress and Failure Conditions 2-8 Cyclic Shear Strength and Normal Stress from Stress-Controlled Dynamic Triaxial Test 2-9 Computed Factor of Safety for the Finite Element Model of UHS Dam 2-10 Undrained Static Strength After Dynamically Loading the Sample (c*)
2-11 Soil Parameters for Stability Anr. lysis - ESWS Pumphouse Channel and UHS Slopes 2-12 Results of Slope Stability Analysis for UHS Excavated Slope Using Wedge Analysis 2-13 Results of Slope Stability Analysis for ESWS Intake Channel 2-14 Results of Consolidation Tests on Undisturbed and Recompacted Soil Samples 2-15 Results of Field Permeability Tests - Ultimate Heat Sink 2-16 Results of Laboratory Falling Head Permeability Tests on Undisturbed and Recompacted Soil Samples 2-17 Monitoring Schedule for UHS and UHS Dam 3-1 Loading and Symbol Descriptions vi SL-3831
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LIST OF EXHIBITS 04-03-81 1-1 General Arrangement 2-1 Ultimate IIcat Sink Plan 2-2 UIIS Elevation-Area-Capacity (as per Record Cross Sections) 2-3 UllS Elevation-Area-Capacity (Estimated) after 40 Years of Sedimentation 2-4 Typical Man-Made Slopes - Ultimate IIcat Sink 2-5 Ultimate IIcat Sink - Intake Channel Cross Section 2-6 Soil Boring and Test Pit Location Plan - UIIS 2-7 Results of Stability Analysis - UllS Dam 2-8 Safe Shutdown Earthquake - llorizontal Ground Spectra 2-9 Safe Shutdown Earthquake - Vertical Ground Spectra 2-10 Artificial Accelerogram for fiorizontal Ground Motion 2-11 Artificial Accelerogram for Vertical Ground Motion 2-12 llorizontal Design Response Spectra for 0.12 g Ilorizontal Ground Acceleration 2-13 Vertical Design Response Spectra for 0.12 g florizontal Ground Acceleration 2-14 Submerged UIIS Dam, Finite Element Representation for Dynamic Analysis 2-15 Cyclic Shear Strength for 5% Strain and 5 Cycles versus Normal Effective Stress 2-16 Dynamic Analysis of Soil Stability Along the Base of UllSD (elevation 1952 feet) 2-17 Results of Static Triaxial Tests After Cyclic Stressing 2-18 Slope Stability Analysis for UliS Dam - Static Strength Following Cyclic Loading I 2-19 Wedge Analysis of Excavated Slopes - Ultimate IIeat Sink 2-20 Slope Stability Analysis, 3:1 Slope, ESWS Intake Channel -
Steady State Condition 2-21 Slope Stability Analysis, 3:1 Slope, ESWS Intake Channel -
Steady State with SSE vii SL-3831 I
04-03-81 2-22 Slope Stability Analysis, 3:1 Slope, ESWS Intake Channel -
End of Construction with SSE 2-23 Slope Stability Analysis, 3:1 Slope, ESWS Intake Chant.el -
Rapid Drawdown Condition 2-24 Slope Stability Analysis, 5:1 Slope, ESWS Intake Channel -
Steady State Condition 2-25 Slope Stability Analysis, 5:1 Slope, ESWS Intake Channel -
Steady State with SSE 2-26 Slope Stability Analysis, 5:1 Slope, ESWS Intake Channel -
End of Construction 2-27 Slope Stability Analysis, 5:1 Slope, ESWS Intake Channel -
End of Construction with SSE
, 2-28 Location of Sediment Alonitoring Pads -
- Ultimate lleat Sink and ESWS Intake Channel 3-1 PAIF flydrographs ( After Construction of Wolf Creek Dam) 3-2 100-Year and Standard Flood Ilydrographs
( After Construction of Wolf Creek Dam) 3-3 Service Spillway Plans 3-4 Tailwater Rating Curve - Wolf Creek Dam 3-5 Service Spillway Wier 3-6 Wingwall Sections Near Weir 3-7 Wingwall Loading 3-8 Wingwall Sections Near Channel Outlet 3-9 Service Spillway 4-1 Auxiliary Spillway Plans
SARGENT & LUNDY ENGIN E E R$
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ENGINEERING DATA COMPILATION FOR WATER-CONTROL STRUCTURES AT WOLF CREEK LAKE WOLF CREEK GENERATING STATION KANSAS GAS & ELECTRIC COMPANY AND KANSAS CITY POWER & LIGHT COM PANY
1.0 INTRODUCTION
1.1 Purpose u of the Report NRC Regulatory Guide 1.127 applies to water-control structures (e.g., dams, reser-voirs, conveyance facilities, etc.) specifically built for use in conjunction with a nuclear power plant and whose failure could cause radiological consequences ad-versely affecting public health and safety. The guide requires that inservice inspec-tions be made periodically to check the condition of the water-control structures and to evaluate their structural safety and operational adequacy. These inspections will include, but will not be limited to, a review of all engineering data related to the design, construction, and operation of the water-control structures.
The purpose of this report is to comply with Regulatory Guide 1.127 requirements by compiling and organizing all engineering data related to the Wolf Creek ultimate heat sink (UllS) dam, UIIS, and main dam service and auxiliary spillways.
1.2 Purpose of the Project The Wolf Creek cooling lake and the associated structures are built to supply cooling water to the Wolf Creek Generating Station (WCGS). The lake is formed by a main dam across Wolf Creek and five saddle dams along the periphery of the lake. The heated discharge from the plant is cooled during circulation in the cooling lake. A long cooling path is provided by baffle dikes and cooling water channels (Exhibit 1-1).
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2.0 UliS DAM AND UHS 2.1 Hydrologic and Hydraulic Design 2.1.1 Design Criteria The ultimate heat sink (UHS) is a submerged pond within the cooling lake which provides sufficient capacity for the essential service water system (ESWS). The design life of the UHS is 40 years.
2.1.2 Design and Analysis The UHS is formed by a Category I dam in one of the fingers of the Wolf Creek cooling lake, it is designed to provide sufficient cooling water for the safe shutdown of the plant. The design is such that the minimum capacity will be available even after the estimated sedimentation occurs in the UHS during its design life.
2.1.2.1 Capacity of the UHS The UHS has a capacity of approximately 455 acre-feet and a surface area of 100.7 acres at a design pool elevation of 1970 feet (SNUPPS Datum).* This elevation is 15 feet and 6 feet below the estimated low water levels of the cooling lake, respectively, for one-unit and two-unit operations of the WCGS. The design bottom elevation of the UHS is 1965 feet (Exhibit 2-1).
- A plant coordinate system, called SNUPPS, is used. The following is the conversion l from SNUPPS to USGS:
SNUPPS elevation = USGS (MSL) elevation + 900 feet SNUPPS E 100,000 = E 2,807,250 of State Plane Coordinate System SNUPPS N 100,000 = N 584,670 of State Plane Coordinate System E 100,000, N 100,000 is the center of the reactor for Unit 1.
The above have feet as units.
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SARGENT & LUNDY W E NGIN E E R S ancaco 04-03-81 2.1.2.2 Sedimentation of the UllS The sedimentation rate for the UllS was estimated assuming that the drainage area contributes runoff into the lake portion east of the UliS dam. The maximum sedi-2 mentation rate is 0.78 acre-feet /mi /yr, which is about 33 acre-feet over a period of 40 years (References 1 and 2). This sediment volume is less than 1% of the storage volume of the lake portion containing the UllS. The UllS will have sufficient surface area and capacity for the essential service water system to function properly after 40 years of sedimentation.
Exhibits 2.2 and 2.3 show the Elevation-Area-Capacity for the UllS for the condition without sedimentation and for the condition after 40 years of accumu'ated sedimentation. The initial Elevation-Area-Capacity was obtained from cecord cross sections of the UllS and ESWS Intake Channel excavation.
2.1.3 Reference Drawings Sargent & Lundy Drawings:
S-80 UllS - Plan and Sections S-81 UliS - Dam - Plan Profile and Sections 2.2 Geotechnical Design 2.2.1 Design Criteria 2.2.1.1 Dam Geometry 2.2.1.1.1 UIIS Dam The embankment is a homogeneous compacted earth (clay) fill. The following design criteria is used:
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- a. The slopes of the dam are designed to withstand all possible conditions l
influencing its stability. The minimum safety factors required under various conditions (Table 2-1) are equal to or higher than those specified in Proposed Guidelines for Safety Inspection of Dams (Reference 3).
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- b. The width of the dam is designed to provide a %fe percolation gradient through the embankment at a full UHS reservoir level at elevation 1970 feet.
I c. The crest width is selected to meet the recommendations of the United States Bureau of Reclamation (USBR) for the design of small dams (Reference 9).
- d. A camber is provided along the crest of the dam to insure that the crest remains at or above its design elevation of 1970 feet after the embankment material has settled.
- e. During a hypothetical main dam failure, which will create an over-topping condition of the UHS dam, riprap and filter materials will protect the entire embankment from scour and embankment erosion.
2.2.1.1.2 UHS Slopes 2.2.1.1.2.1 Slope Characteristics The UHS shown schematically in Exhibit 2-1 consists of a 95-acre Category I sub-merged pond with a design depth of 5 feet. It is formed by excavating and damming a portion of the 5090-acre main cooling lake. The maximum design elevation of the cooling lake is 1995 feet, and the normal operating level is 1987 feet. The full ultimate heat sink reservoir level and the dam crest elevation are both 1970 feet.
The pond floor elevation is 1965 feet. The soil blanket above the bedrock in the UHS area averages 6 feet thick and ranges from 0.3 foot near the UHS dam to a maximum of about 18 feet near the ESWS pumphouse.
2.2.1.1.2.2 Description of UHS Slopes The natural slopes surrounding the UHS pond are very flat, ranging from 1 vertical to 15 horizontal to i vertical to 60 horizontal. The rolling topography gradually slopes toward Wolf Creek and varies in elevation from about 1987 feet at the edge of the cooling lake to 1954 feet at the UHS dam. The topography of the natural grouno and the limits of cuts and fills are shown in Exhibit 2-1.
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a nc.v o 04-03-81 I The man-made slopes forming the periphery of the ultimate heat sink are excavated from elevation 1970 feet to elevation 1965 feet at grades varying from 1.0% to 6.7%
The slopes between the existing grade elevation and elevation 1970 feet are designed to be 5 horizontal to 1 vertical. Exhibit 2-1 shows the plan view of the locations of I these slope cross sections. Exhibit 2-4 presents details of the typical man-made slopes forming the ultimate heat sink.
The essential service water intake channel is located in the northwestern portion of the UHS. Details of the intake channel slopes are presented in Exhibits 2-1 and 2-5.
The excavated slopes in the channel consist of a 5:1 slope from the existing grade to elevation 1970 feet and a 3:1 slope from elevation 1970 feet to elevation 1965 feet.
A 55-foot bench is provided at elevation 1970 feet along the intake channel to protect against blockage by sheet ice. The benching details are shown in Exhibit 2-5.
The ESWS discharge structure is in the eastern arm of the ultimate heat sink. The excavated slopes near the discharge structure are very flat, having a maximum slope of 1% below elevation 1970 feet. The slopes between the existing grade and elevation 1970 feet are 5 horizontal to 1 vertical. A typical cross section of the slope at this location is shown in Exhibit 2-4.
2.2.1.2 Soils The soils for structural fill to construct the UliS dam will be from onsite excavation within the UlIS reservoir and will be selected to provide the most impervious clayey material. Laboratory tests were conducted on soil samples obtained from borings and test pits as shown in Exhibit 2-6. These tests have shown that the material from within the UllS reservoir is suitable for embankment construction.
2.2.1.3 Filter Material l
The fine filter material is a mixture of natural river sand blended with limestone
, chips. The coarse bedding is obtained by crushing Plattsmouth limestone from the Wolf Creek onsite quarry. The design gradation is such that a) no significant head is lost in flow through the filters, and b) no significant invasion of soil is permitted into the filter. The filter materials satisfy the quality criteria for concrete aggregate as specified in ASTM C-33.
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SARGENT& LUNDY ENGINEl R5 a ncac 04-03-81 2.2.1.3.1 Gradation The gradation requirements of the filter material are based on particle size aclation-ships which were developed by Terzaghi and later extended by the Corps of Engineers (References 4 and 5). The UHS dam filter material is designed to meet the following criteria:
I D(15) Riprap < 10 D(85) Coarse Filter 4 < D(15) Riprap 20(1)
D(15) Coarse Filter D(50) Riprap < 25 D(50) Coarse Filter D(15) Coarse Filter < 5 D(85) Fine Filter 4 < D(15) Coarse Filter < 20(1)
D(15) Fine Filter D(50) Coarse Filter < 25 D(50) Fine Filter D(15) Fine Filter < 5(2)
D(85) Soil 4< D(15) Fine Filter < 20(1,2)
I, D(15) Soil D(50) Fine Filter 25(2)
I D(50) Soil
'I (1) This limit may be increased to 40 if the finer material is well-graded (uniformity coefficient: D 60/D10 ' 4)*
(2) These criteria need not be satisfied if the resulting filter material contains more than 5% fines ( < 0.074 mm, No. 200 sieve).
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where D(15), D(50), and D(85) are the particle sizes from a particle size distribution plot at 15%, 50%, and 85x 'iner by weight, respectively. The gradation relationship between the filter and the riprap layer was designed using the Corps of Engineers criteria (Reference 6) for which the D(15) size of the riprap does not exceed 10 times the D(85) size of the filter.
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2.2.1.3.2 Thickness The thickness of the filter is selected based on the following factors:
- a. the wave action of the lake waters,
- b. the gradation of the riprap, and
- c. plasticity and gradation of the embankment materials.
2.2.1.4 Riprap Material During the unlikely postulated total loss of the main cooling lake dam (Main Dam),
the slopes and crest of the UIIS dam will be subjected to a flow of water over the crest. Therefore, adequate erosion protection is required for the upstream and down-stream slopes, as well as for the crest of the UllS dam. The riprap materials satisfy the rock quality requirements for concrete aggregates es specified in ASTM C-33, 2.2.1.4.1 Gradation The riprap is a well-graded material in accordance with a) the guidelines established in the Corps of Engineers publication entitled " Stability of Riprap and Discharge Characteristics, Overflow Embankments, Arkansas River, Arkansas" (Reference 7);
and b) the design techniques presented by Olivier (Reference 8).
2.2.1.4.2 Thickness The thickness of the riprap material is conservatively selected based on the results of the investigations conducted by the Corps of Engineers for similar structures subjected to similar field conditions (Reference 7).
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SARGENT & LUNDY E N GIN E E R 5 a ncaco 04-03-81 2.2.1.5 Compaction The cohesive fill materials are compacted to at least 95% of the maximum dry den-sity, with moisture contents within +2% of the optimum moisture content, as deter-mined from the Standard Proctor Test, ASTM D-698-70. The soil fill materials are placed .'n uniform lifts with loose lift thickness not exceeding 8 inches for heavy compaction equipment or 3 inches for hand-operated power tampers or similar special equipment. The granular material for the bedding course beneath the riprap has a maximum loose lift thickness of 18 inches and is compacted to 80% relative density.
The relative densities are as determined by ASTM D-2049. No compaction requirement is specified for the dumped riprap material.
2.2.1.6 Safety Factors - Stability Analysis The slopes of the dam are designed to withstand all possible design conditions in-fluencing their stability. The minimum safety factors required under various con-ditions (Table 2-1) are equal to or higher than those specified in Proposed Guidelines for Safety Inspection of Dams (Reference 3).
2.2.2 Geotechnical Data 2.2.2.1 Soil and Rock Investigation 2.2.2.1.1 %bsurface Exploration The foundation conditions underlying the water-control structures were investigated by detailed geologic studies, a number of borings, test pits, and geophysical surveys.
The results of the subsurface investigation, which was performed by Dames & Moore (D&M), can be found in the following list of "Geotechnical Investigation" reports prepared by Dames & Moore.
Report Date of Report
- a. Proposed Cooling Lake February 19,1976
- b. On-Site Rock Quarry Areas June 9,1976
- c. Soil Borrow Materials July 6,1976 SL-3831
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- d. Category I Pond & Dam Ultimate Heat Sink October 15,1976
- e. ESWS Pipelines, Pumphouse, &
Discharge Structures August 12,1977
- f. Proposed On-Site Toronto Rock Quarry July 21,1977 In general, the subsurface soils encountered are clayey in nature and are underlaid by shallow rock. The rock is composed of interbedded layers of shale, siltstone, and limestone.
2.2.2.1.2 Testing Dames & Moore tested the soils which were derived from the borrow areas and which are to be used as embankment materials. The following tests were conducted on borrow soils from various parts of the UHS:
- a. Moisture Content
- b. Atterberg Limits
- c. Particle Size Analysis
- d. Specific Gravity a
- e. Compaction
- f. Triaxial Compression Tests (1) Consolidated Undrained (CU)
(2) Unconsolidated Undrained (UU)
- g. Swell Test
- h. Consolidation l
l i. Permeability J. Dispersive Soil Test
- k. Resonant Column Test
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SARGENT & LUN DY W E N GIN E E R 5 ancaco 04-03-81 I The description and results of these tests are given in the Dames & Moore reports listed in Subsection 2.2.2.1.1.
2.2.2.2 Seismology The seismological data are provided in Section 2.5.2 of the Wolf Creek FSAR.
2.2.3 Design Analysis 2.2.3.1 Filter and Riprap Materials 2.2.3.1.1 General The riprap is obtained by blending the Plattsmcuth Limestone from the onsite querry with large size rock from the Southbend Limestone from the Fogle Quarry in Ottawa, Kansas. Tests on these formations were conducted by the Kansas State Highway Commission in Coffey County, and by Dames & Moore. The results of the tests conducted by the Kansas State liighway Commission are summarized in Table 2-2; the results of the Dames & Moore tests are given in the reports listed in Subsection 2.2.2.1.1.
2.2.3.1.2 Materials for Ul!S Dam During the unlikely postulated total loss of the main cooling lake dam and baffle dike 'A', the slopes and crest of the UllS dam will be subjected to a flow of water over the crest. Adequate crosion protection has been provided for the upstream and downstream slopes, as well as for the crest of the dam. The techniques for the design of rock sections for overtopping were presented by Olivier (Reference 8); a series of laboratory tests were made with various sizes of stones to develop paramoters for different flow rates. The test results were applied to the design of the UHS dam slope protection. Following the criteria as described in Subsections 2.2.1.3 and 2.2.1.4 for filter and riprap materials, and in accordance with the guidelines established in the Corps of Engineers publication entitled " Stability of Riprap and Discharge Characteristics, Overflow Embankments, Arkansas River, Arkansas" (Reference 7), the riprap will be a well graded material with the following gradations:
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Maximum Size Weight: 3200 pounds 85% Size Weight: 1500 to 2200 pounds ;
50% Size Weight: 190 to 400 pounds !
15% Size Weight: 25 to 50 pounds Minimum Size Weight: 5 pounds The averrge of a number of gradation tests should fall in the middle of the band given.
The basic criteria or conditions for the UHS dam are quite similar to those expe-rienced and investigated in the Corps' publication. The side slopes used in their study, 4 horizontal to i vertical, are the same as those for the UHS dam. The duration of the overtopping is approximately equalin both cases. The gradation for the riprap for the UHS dam was made to compare to the A-gradation used by the Corps. The UHS dam riprap is twice as thick as that used by the Corps (4 feet as opposed to 2 feet),
and is complemented by two 18-inch filters consisting of a fine filter and a coarse filter. The maximum average water velocity expected over the UHS dam is less than 10 fps, while the Corps had experienced velocities as high es 13 fps.
By examining various flow conditions over the UHS dam which take the tailwater elevation downstream and the headwater elevation upstream 250 feet from the crest of the dam (in contrast to the 100-foot distance used by the Corps), the riprap is found to be in the stable region for nonaccess-type embankments with a gradation of A-1, Plate 48. This A-1 gradation performed similarly to the A-gradation, as described in the Corps' publication (Reference 7).
The riprap material will be 4 feet thick, measured perpendicular to the slopes of the embankment. The filter material (coarse and fine beddings) to be placed under the riprap was designed according to the criteria established in Subsection 2.2.1.3. Based on these criteria, the following gradation sizes are required for the filter material:
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i Coarse Filter Fine Filter Sieve % Passing Sieve % Passing 4 inch 100 3/4 inch 100 3 inch 85-100 1/2 inch 90-100 1-1/2 inch 55-85 3/8 inch 70-100 3/4 inch 30-65 No. 10 20-65 3/8 inch 10-30 No. 30 8-35 No. 4 0-15 No. 50 3-15 No. 10 0-3 No. 200 0-5 Each of the coarse and fine bedding layers will be 18 inches ' hick, measured perpen-dicular to the side slopes. Details of the riprap and filter are shown on Sargent &
Lundy Drawing No. S-81.
2.2.3.1.3 Materials for Slopes Adjacent to ESWS Pumphouse The slopes adjacent to the ESWS pumphouse will be protected by riprap and filter beddings as shown in Sargent & Lundy Drawing No. S-184. The protection consists of 3-foot thick riprap and 3-foot thick underlying filter beddings.
The size requirements for the materials will be the same as those Ihr the main dam.
The design methods are given in report SL-3830, " Engineering Data Compilation for Wolf Creek Lake."
2.2.3.2 Crest Width The crest width of an earthfill dam depends on the following considerations: a) the nature of the embankment materials and the minimum allowable percolation distance through the embankment at normal reservoir water level, b) the height and impor-tance of the structure, c) possible roadway requirements, and d) practicability of construction. A minimum crest width should provide a safe percolation gradient through the embankment at the level of the full reservoir. The crest width was determined by the following procedure suggested by the United States Bureau of Reclamation (USBR) in Design of Small Dams (Reference 9). ,
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SARGENT & LUN DY ENGIN E E R5 cmcaco 04-03-81 W = f + 10 where w = width of the crest in feet z = height of dam in feet above the stream bed (18 feet)
For the UllS dam, a crest width of 20 feet is provided.
2.2.3.3 Slope Stability Analysis 2.2.3.3.1 General As stated earlier in the report, the analysis was performed for conditions affecting the stability of water-control structures.
2.2.3.3.2 Ultimate IIcat Sink Dam (UHSD) Stability 2.2.3.3.2.1 Shear Strength of Materials To evaluate the shear strength of the soil, a series of laboratory triaxial tests was performed on soil samples from within the UIIS reservoir. The soil specimens tested represented t' e range of materials found in the UIIS reservoir limits and adjacent areas. The laborttory test samples were compacted and tested at optimum water content plus 3% with a density of 95% of Standard Proctor (ASTM D-698-70), which was selected to simulate the conditions during construction.
To determine the strength of compacted impervious soils, the following tests were used. Undrained tests were performed primarily to determine the relationship be-tween the shear strength and normal pressure in terms of total stresses for use in the analysis of dam stability during and immediately after construction. Consolidated undrained tests were performed with the pore pressure measured to determine the strength parameters in terms of effective stress (c' and 4'). The purpose of the test was to obtain the strength for use in the effective stress method of analysis. To obtain the lowest shear strengths, the tests were conducted on compacted samples which were completely saturated. The test results are shown in Tables 2-3.1, 2-3.2, and 2-3.3.
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SARGENT & L UN DY M E NGINI E R5 o nc.v o 04-03-81 Rock strengths listed in Table 2-4 are based on laboratory unconfined compressive tests on Nx rock core from the plant and ESWS pumphouse site areas. The shear strength and modulus characteristics of the rock are of such strength relative to the soils that they may be excluded from consideration in the dam stability analysis.
2.2.3.3.2.2 Static Stability Analysis A stability analysis was performed for the most critical section of the dam under the following four conditions:
- a. the end of construction;
- b. rapid drawdown from elevation 1987 feet to elevation 1950 feet;
- c. steady state seepage, reservoir at elevation 1970 feet; and
- d. fully submerged condition, reservoir at elevation 1970 feet.
Tables 2-3.1 through 2-3.3 and Exhibit 2-7 indicate the soil parameters used in the unalysis. For each condition, Exhibit 2-7 also shows the critical circle of failure with the minimum factor of safety. The increase in strength due to filters and large riprap material was not considerea in the analysis. The computer program SLOPE I (Appendix A), which determines the stability of the embankment using the modified Bishop method of slices for a circular failure arc, was used to analyze the various design conditions.
The end-of-construction case was examined using a total stress analysis. The avail-able minimum factor of safety is 2.45, which is higher than the required 1.4.
l The effective stress method of analysis was used to evaluate the rapid drawdown condition. In the analysis, it was assumed that the drawdown was instantaneous and that no drainage occurred while the water level dropped due to the postulated main dam failure. The drained strength was determined by accounting for the stresses to which the soil was consolidated before drawdown. The available minimum factor of safety is 2.18, which is higher than the required 1.2.
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1 I SARGENT & LUN DY E NGIN F ERS
( mcaco 04-03-81 For the steady state seepage condition, an effective stress analysis was used with the pore pressure estimated from a flow net. The shear strengths used were determined from consolidated undrained tests on saturated samples to which pore pressures were applied to simulate those which may exist under the gravity flow of the dam. It is possible that horizontal layers of a material which is relatively more pervious than the average may be built into the impervious section; therefore, a conservative assumption was made concerning the ratio of the permeabilities in the horizontal and vertical directions. The degree of anisotrophy was conservatively selected as 9. The available minimum factor of safety is 2.5, which is higher than the required 1.5.
The fully submerged condition of the embankment was analyzed using an effective stress analysis, with pore pressures corresponding to equilibrium conditions of the main cooling lake (water at 1970 feet or above on either side of the embankment).
The available minimum factor of safety is 4.67, which is higher than the required 1.5.
2.2.3.3.2.3 Seismic Stability Analysis The following two methods of analysis were used to evaluate the seismic stability of the UIIS dam: a) the pseudostatic analysis using the soil strength parameters discussed in Subsection 2.2.3.3.2.2, and b) the finite element analysis using the method proposed by Seed et al. (References 10 and 11).
2.2.3.3.2.3.1 Pseudostatic Analysis in the pseudostatic analysis method, a horizontal force corresponding to the safe shutdown earthquake (SSE) level of 0.12 g was applied to the static conditions analyzed in Subsection 2.2.3.3.2.2. The conditions analyzed were the end of construction, steady state seepage, and fully submerged, with the same soil parameters used for the respective static analyses. The effects of the SSE seismic loading were not applied to the rapid drawdown condition since this would require both events to occur simultaneously, which is not a reasonable assumption. A flow net was constructed to determine the phreatic line for the steady state seepage case.
Exhibit 2-7 shows the critical circles of failure with their minimum factors of safety for the three cases analyzed. In each case, the available minimum safety factor was higher than the required (see Table 2-1).
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sARcLNT& LUN DY E NGINL 0 R5 a nuco 04-03-81 l
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Table 2-5 compares the minimum safety factors obtained against those required for l
various static and pseudostatic conditions analyzed fw the UIIS dam. The UllS dam was found to be safe for all these conditions.
2.2.3.3.2.3.2 Finite Element Analysis 2.2.3.3.2.3.2.1 Introduction In the finite element method of analysis, the procedure used to evaluate the seismic stability of the ultimate heat sink dam consists of the following steps.
- a. A dynamic response analysis of the UIIS dam was conducted to evaluate the shear stress time history at various locations through the embank-ment. The response computation was performed using the finite element method of analysis. The computcr program used to compute the response, QUAD 4, incorporated the strain-dependent modulus and the damping ratio for each element of the model(Appendix A).
- b. The irregular shear stress time histories obtained for the various loca-tions throughout the embankment were represented by equivalent uniform shear stresses corresponding to specified numbers of cycles.
- c. Analysis was performed to determine the static stresses existing in the embankment before the earthquake.
- d. The cyclic shear stresses required to cause strains greater than 5% for conditions representative of those existing in the embankment were determined through appropriate cyclic loading triaxial compression tests on representative specimens of the materials.
- e. The seismic stability of the embankment was evaluated by comparing the shear stress required to cause strains greater than 5% with the equivalent shear stresses induced by the safe shutdown earthquake.
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SARGENT & LUN DY B E NGIN E E R 5 aucmo 04-03-81 The soil properties for the dam correspond to those expected for the constructed dam, and were based on laboratory test results, field measurements, and published and unpublished data.
A comprehensive series of dynamic triaxial compression tests was conducted to evaluate the strength characteristics lnd dynamic properties of the remolded, sat-urated specimens. Strength characteristics of the material were obtained from the stress-controlled dynamic triaxial compression tests. The dynamic properties, shear modulus, and damping ratio were obtained from the strain-controlled dynamic triaxial compression test.
2.2.3.3.2.3.2.2 Design Earthquake and Loading Conditions The safe shutdown earthquake of 0.12 g was considered in the free field at the foundation level of the Category I ultimate heat sink dam. The horizontal and vertical design response spectra for the safe shutdown earthquake (SSE) of 0.12 g horizontal ground acceleration are shown in Exhibits 2-8 and 2-9. In accordance with the design criterion, the dam will remain stable, assuming that the horizontal and vertical accelerations act simultaneously, and while the water level is at the steady state design water surface elevation of 1970 feet.
I 2.2.3.3.2.3.2.3 Procedure Used in Seismic Stability Evaluation The following steps were used to evaluate the seismic stability of the UllS dam,
- a. Generation of Synthetic Accelerograms: The computer program RSG (described in Appendix A) was used to generate synthetic accelerograms for horizontal and vertical motions such that the response spectra of these accelerograms essentially enveloped the design response spectra.
These normalized accelerograms are shown in Exhibits 2-10 and 2-11.
I The close matching of the response spectra obtained for the artificial accelerograms with the design response spectra is demonstrated in Exhibits 2-12 and 2-13.
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l SARGENT & LUN DY E N GIN E E R5 cmcaco 04-03-81 I b. Dynamic Response Analysis: The horizontal and vertical rock motions 1
obtained in Step a. above, were simultaneously used for the dynamic response analysis of the dam to evaluate the shear stress time histories at various locations in the dam. Results of the response computation provided values of cyclic stress that are likely to be induced in the soils during an earthquake. The response of the dam was obtained using the finite element method of analysis.
Exhibit 2-14 shows the finite element representation for the submerged ultimate heat sink dam. The dynamic material properties were incor-porated into the analysis using strain-dependent modulus and damping values. The computer program QI;.6Di (Reference 12) was used to compute the response. This program - described in Appendix A.
Shear modulus and damping values were assigned to each element in the dam. Because these values were strain-dependent, they were not known at the start of the analysis, and an iteration procedure was required. At the outset, values of shear moduli and damping were estimated and the analysis was performed. Using the computed values of average strain l developed in each element, new values of modulus and damping were determined from the appropriate data relating these values to strain.
Proceeding in this way, a solution was obtained incorporating modulus ig E and damping values which were compatible with the average strain developed, and the shear stress time history in each element of the dam was generated.
- c. Representation of Irregular Shear Stress Time History by Equivalent Uniform Shear Stress: The procedure used to represent the irregular shear stress time history of any element by an equivalent uniform shear stress corresponding to any number of cycles is similar to the method proposed by Lee and Chan (Reference 13).
I
- d. Static Stress Analysis: A knowledge of the initial static effective stress conditions was required to evaluate the cyclic strengtn of materials SL-3831 l
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SARGLN f A LUN DY I: N GI N I' ! R 5 cincmo 04-03-81 in the dam. For this purpose, an incremental finite element approach that simulated the construction of an embankment in a series of layers was used. The dam was divided into several horizontal layers, each represented by quadrilateral elements. During any increment of the layer, sppropriate values of the Young's modulus E and Poisson's ra-tio vwere assigned to each element. After determining the stresses, E and u were reevaluated for the average stress conditions during the new increment und were compared with the assigned values. If a significant difference was obtained, the E and u values were adjusted until a reasonable correspondence was established between the input and the computed values. This process contir.ued until the last layer was added. The effect of buoyancy on stresses was evaluated by using the submerged unit weight for the matecial in the dam. The analysis was conducted using the computer program ISHILD, which is described in Appendix A. Table 2-6 provides the soil properties used in the static stress analysis of the UllS dam. The finite element model for this analysis was similar to the one used for the dynamic analysis,
- c. Dynamic Material Prcperties: For this analysis, it was necessary to determine the cyclic shear stresses required to cause strains greater than 5% in the material of the dam for conditions representative of those existing in the dam before earthquake loading. These data were I obtained by appropriate cyclic loading triaxial compression tests on representative specimens of the material conducted in accordance with the procedures described by Seed and Peacock (Reference 14).
- f. Evaluation of Seismic Stability: The initial stress conditions and the failure conditions for the dynamic triaxial test specimens are given in Table 2-7. Following the computation procedure suggested by Seed, et al. (Reference 11), the cyclic load test data in the table led to the results presented in Table 2-8. These results are plotted in Exhibit 2-15. In this figure, the initial stress conditions on a soil element are expressed by the following values: fe, the normal stress on the potential failure surface before the carthquake; T gg, the shear stress SL-3831 I
SARGENT & IUNDY E NGIN E E R5 cmcaco 04-03-81 on the same surface at the same tisi:e; and a fe = r l fe. For different initial stress conditions, the figure shows the values of cyclic shear strength to be applied in the direction of potential failure to cause 5%
axial strain in five cycles. The laboratory test data provides results for values of a equal to 0.108 and 0.288. The stre.cs conditions causing 5%
axial strain for other values of a have been interpolated and plotted in Exhibit 2-15.
The initial static normal and shear stresses were computed along sev-eral planes within the UllS dam. Typical values of initial effective normal stress, a ' , initial shear stress, T 9, and the ratio r g/ o g' along the base of the UHS dam are presented in Exhibit 2-16. These values and the cyclic test data presented in Exhibit 2-15 were used to determine the cyclic strength required to cause 5% strain in five cycles.
The minimum factors of safety for various elements against local failure due to seismic loading were determined by comparing the shear stresses required to cause strains greater than 5% with the equivalent shear stresses induced by the simultaneous action of horizontal and vertical rock accelerations. The induced equivalent uniform shear stresses were determined using ae procedure described in Step c.
above.
In the finite element dynamic stability analysis, seismic stability was evaluated by comparing the shear stress r g required to cause 5% strain at any location to the shear stress rd induced by the safe shutdown earthquake. The ratio was considered to represent a local factory of safety against the development of 5% strain. Based on previous experience (Reference 11), a minimum value of stress ratio r g/ T d greater than 1.1 is considered to provide ample margin of safety for seismic stability. Table 2-9 gives the computed factor of safety for the finite element model of the submerged UHS dam. Since the model is symmetric, the factors of safety are provided in the table for only half of the model elements. The rninimum safety factor is 1.35, which is higher than the required 1.1.
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SARGENT A LUNDY ENGIN0LR$
o"C^'
04-03-81 The results of the dynamic analysis indicate that the UllS dam will have ample margin of safety under the seismic loading conditions. l 2.2.3.3.2.4 Stability Analysis Using Static Strength Following Cyclic Loading To assess the effects of cyclic straining on the soil strength, undrained static loading tests were performed on the compacted samples and on the samples which had been subjected to cyclic loading. The results of the static triaxial tests on samples which were subjected to cyclic loading are shown in Exhibit 2-17. The loss of strength due to 11 cycles was determined by the ratio of the undrained strength of the sample after 11 cycles to the static undrained strength. The test results for loss of strength (Table 2-10) indicate that for"3e = 600 psf and K g= 1.0, the sample retains undrained shear strength e of 390 psf after straining fo,11 cycles. The stability analysis of the UllS dam was performed using the retained undrained shear strength in the computer program SLOPE. The stability analysis for the end-of-construction condition for the dam showed a minimum factor of safety of 2.07 for a e value of 380 psf (which is slightly lower than the c = 390 psf determined by laboratory tests). Exhibit 2-18 shows the critical circle of failure with the minimum factor of safety. This indicates that the use of design parameters following cyclic loading provides an ample margin of safety for the dam.
2.2.3.3.3 Stability of ESWS Facility Slopes I 2.2.3.3.3.1 Natural Slopes TNe natural slopes in the UllS area are quite flat, ranging from i vertical to 15 hori-zontal to I vertical to 60 >rizontal. Failure of the main dam will result in drainage I of the water in the essential cooling pond area, from the normal operating level of 1987 feet to the ultimate heat sink dam crest, elevation 1970 feet. Even though the soils in the natural slopes will be saturated, the residual strength of the natural soil is more than adequate to forestall any slides from occurring on such flat slopes. Any i minor sloughing that may occur at the head of the small valleys will not travel more than a few feet, and will definitely not affect the essential cooling pond storage. Due to the characteristics of the residual soils, the creation of subaqueous flows is not deemed proible. Moreover, there is no evidence of slides having occurred in the area.
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1 sARGENT a LUNDY 04-03-81 I NGINL t R5 t il!CAGO 2.2.3.3.3.2 Man-Made Slopes The profiles of the excavated (man-made) slopes in the UHS, the intake channel, and near the outlet structure are discussed in Subsection 2.2.1.1.2.2. The slope protection required for the slopes adjacent to the pumphouse is discussed in Subsection 2.2.3.1.3.
l 2.2.3.3.3.2.1 Ultimate Heat Sink Slopes The UHS slopes forming the periphery of the ultimate heat sink were designed to be 5 horizontal to I vertical between the existing grade elevation down to elevation 1970 feet. From elevation 1970 feet to 1965 feet, the grades vary from 1.0% to 6.7% Typical cross sections of these slopes are shown in Exhibit 2-4.
The soil cover in the excavated UHS slopes is thinner and weaker than the bedrock existing at shallow depth. A circular failure are would not develop for these slopes; therefore, their stability was investigated using the wedge method of analysis. In this method, it is assumed that if failure should occur, it would be by a sliding-block type mechanism with a vertical boundary between the blocks. The potential failure mass is broken into two or three wedges. The shear resistance along segments of the failure surface is expressed in terms of the applicable strength parameters.
The soil properties used in the stability analyses are shown in Table 2-11. The critical cross section analyzed is shown in r.xhibit 2-19. The following conditions were analyzed:
- a. the end of construction,
- b. steady state seepage, and
- c. steady state seepage plus SSE (0.12 g).
The end-of-construction case was examined using a total stress analysis in which parameters c eu and c cu, e tresponding to the consolidated undrained condition, were used to evaluate the shear strength of the soil. For steady state seepage condition, an effective stress analysis was used.
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1 SARGliNT A !.U N DY tucanuns
< m e u.o 04-03-81 The minimum safety factors computed and those required for various cases, are summarized in Table 2-12. There is an ample margin of safety for all cases con- l sidered.
2.2.3.3.3.2.2 ESWS Pumphouse Intake Channel Slopes The excavated slopes in the channel consist of 5:1 slopes from existing grade down to elevation 1970 feet, and 3:1 slopes from elevat'on 1970 feet to elevation 1965 feet.
There is a 55-foot bench at elevation 1970 feet. The two slopes (upper slope 5:1.and lower slope 3:1) were analyzed for the following conditions:
- a. steady state, water in channel at elevation 1970 feet,
- b. steady state with SSE of 0.12 g,
- c. end of ecnstruction,
- d. end of construction with SSE of 0.12 g, and e, rapid drawdown from elevation 1970 feet to 1965 feet.
The BlSIIOP computer program was used to investigate clope stability for the above design conditions. Details of this program are given in Appendix A. The soil proper-ties used in the analyses are given in Table 2-11. Shale rock was conservatively assumed to have the properties of residual clay shown in Table 2-11 for the stability analyses.
I The effective stress method of analysis was used to evaluate the steady state condi-tion with and without an SSE of 0.12 g. The minimum factors of safety obtained for l the static case were 7.13 for the 3:1 slope and 3.37 for the 5:1 slopes; the minimum factors of safety with SSE effects were 3.37 and 1.86, respectively. These factors of safety are higher than required, as indicated in Table 2-13. The safety factors for the l
flatter slope (5:1) are lower than those for the steeper slope (3:1) because the 5:1 slope is much greater in height than the 3:1 slope.
Consolidated undrained total stress parameters were used to analyze the end-of-construction condition. The minimum factor of safety obtained without SSE effect
~
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SARGLN F A LUNDY tqcmt tns 04-03-81 cmcaco for the 5:1 slope was 3.14. With the SSE cffect, the safety factors obtained were 3.88 and 1.75 for slopes 3:1 and 5:1, respectively. No analysis was performed for the static case of the 3:1 slope since the available factor of safety with the SSE is considerably higher than 1.5.
Consolidated undrained total stress parameters were used to evaluate the rapid drawdown condition. In this analysis, the drawdown was assumed to be instantaneous, with no drainage occurring while the water level dropped. Rapid drawdown in the channel will have no effect on the stability of the 5:1 slope. For the 3:1 slope, the factor of safety obtained was 5.69. Table 2-13 summarizes the computed and required factors of safety for the various cases analyzed. The intake channel slopes were determined to be stable under all conditions considered. Exhibits 2-20 through 2-27 show failure circles with minimum safety factors for all cases analyzed.
2.2.3.3.3.2.3 Slopes Near the Discharge Structure The excavated slopes near the discharge structure are very flat, having a maximum slope of 1% below elevation 1970 feet. The slopes between the existing grade and elevation 1970 feet are the same as those in the UIIS. The stability analyses of thse slopes are described in Subsection 2.2.3.3.3.2.1. Based on the results of these analyses, it is concluded that the excavated slopes of the UllS, the intake channel slopes, and the slopes near the outlet structure are stable under all extreme design loading conditions.
2.2.3.4 Settlement Since the embankment for all water-control structures is founded on rock or very close to rock, most consolidation will occur within the fill material placed. The laboratory consolidation tests were performed on that fraction of the embankment soil which is finer than the No. 200 sieve. The results of consolidation tests for both the compacted and the undisturbed soil samples are shown in Table 2-14.
A camber, as discussed in Subsection 2.2.3.5, is provided along the crest of the UHS dam to insure that the crest of the dam remains at or above its design elevation after the embankment material has settled. To further check the settlement of SL-3831
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SARGENT & !.U N DY I NGIN E f R S c uxxo 04-03-81 water-control structures, settlement points are installed along the centerline of the dam crest. Details of the settlement points are given in Section 2.2.4.
2.2.3.5 Camber Camber is usually provided along the crest of earthfill dams to insure that the free-board will not be diminished by foundation settlement or embankment consolidation.
Selecting the amount of camber is somewhat arbitrary; it is based on the amount of foundation settlement and embankment consolidation expected for a particular dam, with objective of providing enough extra height so that some residual camber will remain after settlement and consolidation. For the UHS dam, most embankment consolidation occurs during construction, before the embankment is completed. The embankment is founded on rock or very close to rock, which is relatively incom-pressible material. Therefore, foundation settlement is small. A camber of 3.5% of the height of embankment is provided. The height of the dam is measured from the rock surface.
2.2.3.6 Seepage Control The degree of seepage control required for the UHS dam was determined by detailed geologic exploration, a number of borings, and test pits at the site. In addition, field permeability tests were made across the entire length of the dam. A series of 150 pressure water-loss tests was made in the total stratigraphic geologic column to a depth of 2.5 times the height of the dam. A summary of the results of the permeability at the dam site, which include the effects of rock jointing, range from 0 ft/yr to 48 ft/yr. A representative value for an upper limit of rock mass permeabil-ity is 100 ft/yr. The conservative upper limit was exceeded by none of the tests conducted at any location within the entire bounds of the ultimate heat sink area. In addition, both falling and constant hee permeameter tests were conducted to complement the pressure test results in specific zones.
The representative permeability value for the natural and recompacted soil samples was obtained from laboratory permeability tests. The test results are shown in
-0 Table 2-16. For compacted embankment clays an average permeability of 3.6x10
-7 cm/see was measured. A value of 1x10 cm/see was used in the seepage analysis to SL-3831 I
SARGENT & !.U N DY W F NGINI I R5 u ncuo 04-03-81 provide a conservative estimate of seepage through the embankment. Based on the average permeability of compacted, well-graded sands and gravels, the permeability
-3 of the filter material was assumed to be 1x10 cm/sec. The rock used as the riprap blanket was assumed to be freedraining. Seepage through the UllS dam was computed using the computer program SEEPAGE (described in Appendix A). Based on the permeability values described above, a discharge 010.23x10-4 efs/ft through the dam was calculated.
Based on the results of the seepage analyses, it was concluded that no special seepage control measures were required. Adequate drainage is provided near the toe of the dam so that softening of the toe does not occur.
2.2.3.7 Solution and Weathering Known solutMn features in eastern Kansas are confined to areas containing thick outdrops c f water-soluble rocks, local buildups of reefoid carbonates, faulting, or stream channel diversions. Since none of these conditions is present at the Wolf Creek site, the possibility of instability due to collapse of solution cavities is con-sidered minimal. The extent of solutioning and weathering of the limestone can be related to the amount of calcium in the water, the amount of calcium that can be retained in solution by the water, and the amount of calcium present in the limestone.
There is evidence of solutioning in the lake work area, but it is of a minor nature.
Investigations to demonstrate the chemical balance between the water and the lime-
! stone sie presented in Section 2.5.1 of the FSAR. Since the lake waters are normally near or above the saturation levels with respect to calcium, no solutioning is ex-pected. In addition, the relatively low permeabilities of the limestones at the site and the low permeabilities of the overlying soil and of the interbedded shales preclude the development of Karst features at the site.
The bedrock units at the site have been weathered to depths ranging from 5 to l approximately 30 feet below the ground surface. Because the rock units are characterized by low permeabilities and because the water table is near the surface, weathering does not extend to greater depths.
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SARGEN1'a 1.UN DY W I NGIN L1 R5 c mcam 04-03-81 Type and degree of weathering are largely dependent upon bedrock lithology. Shale units are typically weathered more deeply and severely than limestone. Weathering of the shales results in discoloration and a decrease in rock strength and consistency; weathering is usually concentrated along bedding planes. The clay minerals com-prising the shales are degraded and hydrated by the weathering process, resulting in a claycy shale matrix.
Limestone units are typically weathered along fracture surfaces and shaley bedding plancs. The limestone acts as a barrier to the weathering process, as weathering extends through a limestone unit only when it is very thin or when it is exposed at or very near the ground surface. At the site, weathering did not extend below the uppermost limestone when the soil cover was 10 feet or more.
2.2.3.8 Liquefaction Potential The UllS dam is designed for the safe shutdown earthquake conoition. Surrounding the UIIS is a very gently rolling area consisting mainly of clayey soils which do not possess any of the characteristics requisite for liquefaction. Granular materials used in construction will be compacted to relative densities of 80% or higher to provide an adequate margin of safety against liquefaction under site seismic conditions.
Therefore, liquefaction is not a problem.
2.2.4 Instrumentation and Monitoring Program The UllS dam and pond will be monitored periodically to assure that the UllS main-tains the required storage capacity. The program will consist of monitoring settle-ment points on the UIIS dam and sedimentation in the UllS reservoir and intake channel. The schedule for making these measurements is given in Table 2-17.
2.2.4.1 Settlement Points on UllS Dam Settlement points will be installed alor.g the axis of the UIIS dam at stations -2+00, 0 + 00, 2 + 0 0, 4 + 0 0, 5 + 5 0, 7 + 00, 8 + 5 0, 10 + 0 0, and 12 + 0 0. They will be mor tored for vertical and horizontal movement of the dam crest, and will consist of 3-foot di-ameter concrete pillars with a survey plug on top, extending 5 feet into the dam I embankment and extending upward through the filters and riprap. The settlement j SL-3831 l
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SM GENT & LUN DY E NGIN E E R 5 cmcaco 04-03-81 point at station -2+00 is in the dam abutment and can be used as a baseline mea-surement point. The locations of settlement monitoring points are shown on Sargent & Lundy Drawing No. S-81.
2.2.4.2 Sedimentation Monitoring Points in UllS Pond and Intake Channel Sedimentation monitoring pads will be installed at 18 locations within the UHS pond and intake channel, as shown in Exhibit 2-28. These will consist of 4 foot by 4 foot concreta pads with their surfaces flush with the UHS bottom surface. The pads will be mark'ed with a rod extending above the pad surface and will allow progressive measurements of sedimentation thickness to be taken at the same loestions.
2.2.4.3 Profile Profiles of the UllS pond and intake channel will be made to determine that the UllS maintains the required volumetric capacity.
2.2.4.4 Visual Observations Visual observations will be made of the area around the UIIS pond, intake channel, and ESWS pumphouse to determine any conditions that may affect the performance of the ESWS system - i.e., decreased design capacity or blocked flow of water. These inspections will occur monthly during lake filling and quarterly thereafter. Inspec-tions will also be made after violent wind storms, flood conditions on the lake, and drawdown of the lake. If it is evident that the potential for intake channel blockage or obstruction may exist, a diver will be dispatched to inspect the intake channel.
2.2.5 Reference Drawings Sargent & Lundy Drawings:
S-81 UllS Dam: Plan, Profile, and Sections S-184 Earthwork Plan & Sections, ESWS Pumphouse SL-3831
SARGENT & LUN DY M
ENGINEERS anc^co 04-03-81 3.0 SERVICE SPILLWAY 3.1 IJydrologic and Ilydraulie Design 3.1.1 Design Criteria
- a. The service spillway is designed in conjunction with the auxiliary spill-way to discharge the probable maximum flood (PMF) with an antecedent standard project flood (SPF). The spillway rating and flood routing program of the U.S. Army Corps of Engineers (Sargent & Lundy Pro-gram SPRAT) is used to route the floods. A description of SPRAT is given in Appendix A.
- b. The peak total outflow from the service and auxiliary spillways under the post-project condition will not exceed the peak flow under the pre-project condition at the dam site.
A c. The spillway design flood (SDF) is the PMF, which is used to determine the maximum water level in the lake. The hydraulic design of the spillway components is done for the SPF, and their structural integrity is maintained while passing the PMF, 3.1.2 Design and Analysis The service spillway is located on the east abutment of the main dam. With the auxiliary spillway, it is designed to safely pass the PMF with antecedent SPF de-veloped for the Wolf Creek drainage basin above the dam site. The PMF and SPF hydrographs are shown in Exhibits 3-1 and 3-2, respectively.
The service spillway is an uncontrolled ogee-crested spillway,100 feet in length and semicircular in plan. Its crest elevation is 1988 feet. A 30-foot wide chute carrit, the spillway discharge to a stilling basin. A pilot channel conveys the flow from the stilling basin downstream to the main channel of Wolf Creek. The ogee crest shape is designed for 75% of the maximum head over the spillway design flood. The chute sidewalls and the stilling basin are designed for the SPF. The chute walls and SL-3831
SARGENT & LUNDY f NGIN L E R $
cnic^co 04-03-81 stilling basin walls are designed to pass the SPF with freeboard and to contain the PM F.
The procedures r"' tined in publications by the Bureau of Reclamation and the U.S.
Army Corps of Engineers (References 9 and 15) are used to design the service spill-way. Details of the service spillway crest, the chute, and the stilling basin are shown in Exhibit 3-3.
- a. Crest Shape of the Spillway:
Water level due to the spillway design flood in the cooling lake = 1995.0 feet Spillway crest elevation = 1988.0 feet Design head (75% of head due to spillway design flood) = 5.5 feet (approx. )
Upstream height of spillway = 13.5 feet Elevation of toe of spillway = 1974.0 feet Design outflow discharge = 5060 cfs The design shape of the ogee crest is shown in Exhibit 3-3.
- b. Design of Chute:
Spillway design flood outflow discharge = 7320 efs Design discharge corresponding to standard project flood = 2700 cfs Width of chute = 30 feet The alignment of the chute is as shown in Exhibit 3-3.
The water surface profiles are calculated using the U.S. Army Corps of Engineers program, " Water Surface Profiles" (Sargent & Lundy program W ASP 7 7). A description of WASP 77 is given in Appendix A. The freeboard is calculated using the procedure outlined in Reference 15.
The top of the chute walls is at a level corresponding to the water SL-3831
I SARGENT & LUNDY l NGIN L f R5 crucaco 04-03-81 surface elevation due to the design discharge plus freeboard or the water surface elevation due to the spillway design flood, whichever is higher.
- c. Stilling Basin:
The tailwakr rating was developed for Wolf Creek and is presented in Exhibit 3-4. A backwater computation is done along the tributary creek to develop the tailwater rating curve (Exhibit 3-4) just downstream of the stilling basin. A stilling basin (Type II) is designed as per procedures outlined in " Design of Small Dams"(Reference 9). The side walls of the stilling basin are designed for the design discharge plus freeboard. The side walls are raised, wherever necessary, to contain the discharge due to the SDF.
Design discharge (outflow due to the standard project flood) = 2700 efs Spillway design flood outflow discharge = 7320 efs Width of the basin = 30 feet The design details are shown in Exhibit 3-3.
3.1.3 Reference Drawings Sargent & Luady Drawings:
S-57 Service and Auxiliary Spillways - Plan S-60 Service Spillway Discharge Channel- Plan, Profile, & Sections S-63 Excavation Plan - Service Spillway S-64 Excavation Sections - Service Spillway S-65 Grade Plan - Service Spillway S-66 Grading Sections - Service Spillway S-460 Service Spillway Plan and Sections - Sheet 1 S-461 Service Spillway Plar, and Sections - Sheet 2 S-462 Service Spillway Plan and Sections - Sheet 3 S-463 Service Spillway Sections and Details SL-3831
SARGENT & LUN DY E N GIN E E R S accaco 04-03-81 3.2 Structural Design 3.2.1 Design Criteria The service spillway structure is designed to allow normal lake overflow at a pre-determined level while preventing erosion or other possible damage. The physical parameters required for proper hydraulic performance of the structure have been discussed in Section 3.1.
Structurally, the service spillway is designed to withstand the effects of water pres-sure, lateral soil pressure, soil weight, self weight, lateral and vertical surcharge and/or 1120 truck loading as described in Table 3-1. For design purposes, the service spillway is divided into five basic components: spillway (weir), wingwalls near the weir, the chute, the road over the channel, and wingwalls at the chute outlet. The chute and wingwall components were analyzed at various locations and designed accordingly to achieve an economy of materials. All components were analyzed by standard, accepted methods, and designed in accordance with appropriate chapters of
" Building Code Requirements for Reinforced Concrete (ACI 318-71)" and " Standard Specifications for Ifighways and Bridges (AASHTO) 1973."
3.2.2 Materials and Allowable Stresses The service spillway is a reinforced concrete structure composed of 4,000 psi mini-mum compressive strength (at 28 days) concrete and ASTM A615 60,000 psi minimum yield strength deformed steel reinforcement.
3.2.3 Design and Analysis 3.2.3.1 Weir The spillway weir shown in Exhibit 3-5 has been designed to resist sliding induced by n/2 R total hydrostatic water pressure calcualted to be / / Wp COSO. de. dr =
170.0 kips, where Wp = (Wg)(HVv ATER) /2. The sliding force is adequately resisted by active soil pressure applied to 29.33 feet of wingwall, or (P g3 MHSOIL) ( 9.33)/2 =
563.7 kips. The minimum factor of safety against sliding equals 563.7/170.0 = 3.32.
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SARGENT& LUNDY E N GIN E E R S cmc ^co 04-03-81 I 3.2.3.2 Wingwalls at Weir End 3.2.3.2.1 General Wingwalls are provided in two separate areas of the service spillway: near the weir and at the chute outlet. Both areas are designed as cantilever retaining walls for dam backfill, with each area investigated at three critical sections. The critical sections I for the wingwalls near the weir are indicated in Exhibit 3-6. Each section has been designed to resist sliding, overturning, and internal forces (moments and shears) induced by the effects of the following combinations of loads:
- a. sliding, overturning, and toe moments and shears = lateral soil pressure (PLS) + lateral surcharge pressure (PSC) + concrete weight (WC ) + soil weight (W3) + vertical surcharge (W SC I b.
wall moments and shears = lateral soil pressure (Pgg) + lateral sur-charge pressure (PSC}'
c.
I heel moments and shears = vertical soil weight (WS ) + vertical surcharge (WSC) + c nerete weight (WC*
A summary of all applied loads is indicated in Exhibit 3-7. The load factor used in the analysis for sliding and overturning is 1.0 for all loads indicated in load combination "a" above. A load factor of 1.7 has been applied to all loads in loading combinations for mcments and shears.
3.2.3.2.2 Sliding Check l
1 Adequate resistance to sliding of the three windwall sections in Exhibit 3-6 was provided, assuming that the total lateral force (Fg) applied to the wingwall was opp =ad bv friction at the bottom of the structure and bearing against rock or soil.
The total lateral force Fg = (PSCN "1
- LS Hg) /2. The resisting force F = [(Vol-ume of Concrete)(W g ) + (WS N1 1 3)/2 + (WS )(H 2 }(b 2 } + ( SC}(b 3 )] u + (PLS)
(H2) /2 + (concrete area bearing on rock) (maximum bearing capacity of rock). For Section 1, the height (H 2) f soil above the base slab toe was negligible and assumed to be zero. Shear keys 2 feet 6 inches deep were cut into the bedrock for Sections 1 SL-3831
SARGENT & LUNDY F NGIN E E R$
04-03-81 1
and 2 to sufficiently resist sliding. The oefficient of friction (p) was taken as 0.2 and the bearing capacity of the rock was 15 kips /ft . In all cases, a 1-foot wide strip was used for loading and analysis. To obtain the factor of safety against sliding, the total applied and resisting forces for the three sections are as follows:
SECTION Fg(kips) Fg(kips) FACTOR OF SAFETY (F R L 1 23.01 46.36 2.01 2 17.86 45.25 2.53 3 13.34 27.07 1.80 3.2.3.2.3 Overturning Check The overturning tendency of wingwalls in Exhibit 3-6 is represented by the total overturning moment of the lateral forces taken about the bottom tip of the base slab toe and is expressed by M g = (PLS)(H1 +t) /6 + (PSC)(Hy +t) /2. Resistance to overturning is achieved by the sum of the moments of the applied verticalloads taken about the same point. The resisting moment is found from MR*( C) (v lume of concrete) ( C.G.(wingwall)] + (WSC} 3 .25 + L2+b3 2) + (W3)(H2}(b2) 2+
(W S 1 "1 2 + (W3 XH1} 3 1 [ 1.25 + L2 + 2 3)13-b1)] /2. A summation of the overturning and resisting moments for the three retaining wall sections is as follons:
SECTION Mg (kip-ft) A1R(kip-ft) FACTOR OF SAFETY (M R !o 1 255.2 622.2 2.44 2 173.6 417.77 2.41 3 104.4 282.74 2.71 I 3.2.3.2.4 Section Design for Shears and Moments Internal design moments and shears have been calculated at the base of the wall and in the slab on both sides of the wall (toe and heel). The shear force at the base of the
+
wall results from lateral surcharge and soil pressures, or V,3;g = (1.7) [ (PSC}("1 (P
LS)(H g )2/21. The maximum shear force on the heel slab is Vheel = (1.7) [(WS SL-3831
SARGENT& LUN DY l' N G I N f: I. R 5 a naco 04-03-81 (L ) + (W g S } 1 "I C)(t)(L1 )]. The maximum shear force in the toe slab is cal-culated from the base pressure resulting from the total vertical loading on the wing-wall, inclLding self weight. The trapezoidal pressure loading shown in Exhibit 3-7 is the effect of eccentricity (e) from the overturning and resisting moments on the structure, or e = LT /2 - I(hi R 'O)/ EW ]. The pressures Pg and P2 at each end of the base slab can be found from the expression P = [E //(L T)(1+6e/(Lt )]. The soil pressure P3 at me inkrseedon of me toe slab and me waH is gP @y4 2 2 h*
- shear force in the toe slab at the same point is V Toe = (1.7) [(P y+P 3 2 ] 4 2). The shear stress at each of these sections is v = (1000V)/(120d), where 8 = 0.85.
The design moments are calcualted at the same locations in the wingwalls as the shears. The moment for the wall U is 51 (Wall) = (1.7) g
/6 + (P SC)(H
[(PLS)(H y)3 b2].
+
The moment in the heel slab is 510 (Hec 0 = (1.0 [(WS }("l}(b )1 /2 + (WS C )Ib1 }
(W ~
2+
C 1 1' 2 1 T) } Ibj/6)]. The toe slab is designed for 51 U e = (1.0 [ P 3 2) /2 + (2) (P g-P3 2 6 - (WC)(t)(L 2 ) /2 -(W 3 )(H 2 } 2 b
The design moments and shear stresses expressed in kip-ft and psi, respectiv ~ 7, are summarized below:
WALL HEEL TOE SECTION S1 U U C U ^I U
V U
V C U bI U
V U
VCVU 1 310.6 63.9 1.98 121.2 107.0 1.18 211.6 97.0 1.30 2 213.0 64.7 1.96 51.6 75.3 1.68 114.8 76.6 1.65 3 116.1 56.7 2.23 37.4 80.6 1.55 88.8 96.4 1.31 The shear capacity of concrete (V C
) is calculated from (2) if'c = 126.5 psi. The moment capacity of concrete members is a function of the area of reinforcing steel and the effective depth of the member. The moment capacities for the various sections and mt mbers based on the appropriate parameters and considering a SL-3831
SARGENT & LUNDY E N G'N E E R S ClHCAco 04-03-81 0.9 capacity reduction factor are as follows: .
I WALL HEEL TOE SECTION M '
- C C.M. Mg C.M.
C 1 346.2 1.11 138.4 1.14 260.8 1.23 2 225.0 1.05 105.6 2.0 115.7 1.01 3 116.1 NOTE 1 65.1 1.74 102.2 1.15 NOTE 1 - The total applied moment for this section was calculated as a function of I the resisting moment of the reinforced concrete section, assuming that passive soil pressure is developed to a level sufficient to exactly balance the difference between the active pressure moment and the reinforced concrete resisting moment.
Properties for granular soil backfill were used in the analysis of the wingwalls near the weir.
3.2.3.3 Wingwalls at Stilling Basin End Wingwall sections at the channel outlet indicated in Exhibit 3-8 were designed for the same general equations and loading scheme used for wingwalls near the weir, except that the total shear force and bending moment in the wall were reduced, taking advantage of the active soil pressure on the toe side of the wall. The resisting active pressure shear force is V U = (1.7) (Pgg)(H2 )/2, and the applied resisting moment is 51U = (1.7) (PLS)(}I2 ) /6. Cohesive rather than granular properties were used for the wingwall at the outlet. The following summarizes the analysis and design of the outlet wingwalls:
! SLIDING lI
- SECTION Fg F R
AMOR OF SAFEH 1 33.2 51.0 1.54 2 12.3 14.9 1.21 3 6.1 9.8 1.61 SL-3831
SARGENT & LUN DY W E NGIN E E R 5 cmCAGo 04-03-81 OVERTURNING SECTION M R ACTOR OF SAFEW O
1 183.0 464.6 2.5 2 103.6 219.9 2.12 3 50.5 129.4 2.56 SHEAR WALL HEEL TOE SECTION V U
V C U Y U VCIYU Y U
VNC U 1 88.9 1.44 91,6 1.38 104.0 1.22 2 76.0 1.66 80.0 1.58 84.0 1.51 3 64.23 1.97 51.0 2.48 40.6 3.12 MOMENTS WALL HEEL TOE SECTION M C.M. M C.M. M U C.M.
U C U C C 1 231.4 231.4 NOTE 1 163.0 213.6 1.31 94.9 208.1 2.19 2 132.8 231.4 1.74 85.5 95.7 1.12 65.1 115.9 1.22 3 73.1 217.9 2.98 42.2 74.4 1.76 19.2 70.8 3.7 i
NOTE 1 - The total applied moment for this section was calculated as a function of the resisting moment of the reinforced concrete section, assuming that passive soil pressure is developed to a level sufficiant to exactly balance the difference between the active pressure moment and the reinforced l
concrete resisting moment.
l l
3.2.3.4 Channel The procedure used to design the spillway chute is similar to that used for the wing-walls. Calculations for moments and shears use the same loading and equations as SL-3831
I SARGEN T & LUN DY
- i. N 31 N E l R 5 a cc.4.c.o 04-03-81 I for the wingwalls, except that the chute walls are loaded on the heel side only. The effective toe length (L )2 shown in Exhibit 3-7 is a variable based on the stiffness of the slab and foundation; it never exceeds half the overall chute width. The tip reinforcing for the slab is determined by the most critical moment found where the I heel portion of the slab intersects the wall. Bottom slab reinforcing is determined by the moment in the toe portion of the slab at the wall which results from vertical soil pressure under the slab. This pressure and the pressure for the wingwalls are calculated in the same way, except where the eccentricity (e) is located outside the middle third of the effective slab length. For this case, the pressure load becomes triangular over a length of base equal to 3 [(L T
/2)- e }.
The chute was analyzed and designed in five sections, as indicated in Exhibit 3-9.
Sections 1, 2, 4, and 5 were designed using granular type backfill; Section 3 was designed using cohesive soil (granular soil wa. < 7tually used in construction). A summation of the design forces is as follows:
SiiEAR STRESSES SECTION WALL IIEEL TOE L V Y Y V C/VU V YCYU T U C U U U 1 24.0 82.3 1.54 105.2 1.20 121.3 1.04 2 15.0 50.3 2.51 64.4 1.96 78.4 1.61 3 19.0 60.0 2.11 97.5 1.30 51.0 2.48 4 21.2 61.3 2.06 69.8 1.81 117.9 1.07 24.0 81.1 1.56 55.6 2.28 82.8 1.53 I
5 l
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SARGENT & LUN DY E N GIN E E R5 cmcaco 04-03-81 MOMENTS SECTION WALL HEEL TOE C.M. M 51 U C U C U "C *
- 1 336.1 341.7 1.02 96.7 117.0 1.21 217.3 218.7 1.01 2 123.3 147.2 1.19 30.7 95.7 3.12 105.1 115.7 1.10 3 40.1 46.2 1.15 23.7 57.0 2.41 53.0 54.3 1.03 4 129.2 144.1 1.12 45.9 95.7 2.08 79.3 81.5 1.03 5 192.1 233.7 1.21 38.2 147.2 3.86 259.2 278.7 1.08 3.2.3.5 Bridge The roadway bridge over the spillway was designed to meet AASffTO requirements.
The bridge is assumed to be simply supported between channel walls. Load factors of 1.4 and 1.7 for dead and live loads, respectively, include a 39% impact factor applied to the live load.
The bridge was analyzed for self-weight, assuming 2-foot thick concrete covered with 4 inches of asphalt (future) and 1120-44 truck loading. The design span of the bridge is 31 feet. The unfactored dead load moment was calculated from the expression (W)(L2 )/8, and equaled 42.3 kip-ft for a 1-foot width of bridge. Live load moments were taken from Appendix A of Standard Specifications for flighway Bridges. The live load moment is 256.5 kip-ft and the moment due to impact is 30% of this, or 77.0 kip-f t. The total live load moment is 333.5 kip-ft over the entire width of the bridge, or 28.5 kip-ft for a 1-foot strip. The total factored moment is MU*
107.7 kip-ft. The moment capacity of the concrete slab section of the bridge is 144.6 kip-ft, giving a capacity margin of 1.34.
The shear force due to live load was taken from Appendix A of Standard Specifica-tions for Ilighway Bridges. The total live load plus impact shear force is found from U = (1/2)(W)(L) and is equal to 5.45 kips /ft. The total factored shear force is 14.47 kips /ft, and the shear stress is uv = VU/ [(0)(12)(d)] = 64.5 psi. The allowable shear stress in a concrete slab is 2 if'c = 126.5 psi. The shear capacity margin is l
l SL-3831 l
l
SARGENT & LUN DY LNGINEFR$
a ncxo 04-03-81 1.96. Edge beams were provided according to AASHTO requirements to resist a live load moment of M = (0.10)(P)(C) = 49.6 kip-ft, where P = 16 kips for 1120 truck and L =
31 feet. The dead load moment for a 3 foot 4 inch by 1 foot 4 inch beam is 79 kip-ft.
The total factored moment is 51 U = 195.5 kip-ft. The moment capacity provided by two No. 8 bars is 260.4 kip-ft, with capacity margin of 1.36.
The chute design for Section 1, Exhibit 3-9, envelops the most critical combination of loading and chute geometry of either soil, truck or bridge loading with lateral support at the top of the walls from the bridge, or soil and surcharge loading without wall support from the bridge.
All reinforcing not indicated in this report has been provided to meet minimum re-quirements of temperature, shrinkage, and distribution reinforcing as specified in the applicable codes and specifications.
3.2.4 Reference Drawings Sargent & Lundy Drawings:
S-57 Service and Auxiliary Spillways - Plan S-64 Excavation Sections - Service Spillway S-65 Grade Plan - Service Spillway S-66 Grading Sections - Service Spillway S-460 Service Spillway Plan and Sections - Sheet 1 S-461 Service Spillway Plan and Sections - Sheet 2 S-462 Service Spillway Plan and Sections - Sheet 3 S-463 Service Spillway Sections and Details SL-3831
SARGENT& LUNDY E NGIN E E RS cmc ^co 04-03-81 4.0 AUXILIARY SPILLW AY 4.1 flydrologic and Hydraulie Design 4.1.1 Design Criteria
- a. The auxiliary spillway is designed in ecnjunction with the service spillway to discharge the probable maximum flood (PMF) with an antecedent standard project flood (SPF). The spillway rating and flood routing program of the U.S. Army Corps of Engineers (Sargent & Lundy Program SPRAT, see Appendix A) is used to route the floods.
- b. The peak total outflow from the auxiliary and service spillways under the post-project condition will not exceed the peak flow under the pre-project condition at the dam site.
- c. The snillway design flood is the PMF, which is used to determine the maximum water level in the lake. The hydraulic design of the spillway components is conducted for the SPF, and the structural integrity is maintained while passing the PMF.
4.1.2 Design and Analysis The auxiliary spillway is located on the east abutment of the main dam. With the service spillway, it is designed to safely pass the PMF with antecedent SPF developed for the Wolf Creek drainage basin above the dam site. The PMF and SPF hydrographs are shown inExhibits 3-1 and 3-2, respectively.
The auxiliary spillway is 1500 feet east of the service spillway and is of the open-cut type. The crest length is 500 feet at elevation 1990.5 feet. The discharge over the auxiliary spillway passes through a downstream protected channel to a natural creek leading to the main channel of Wolf Creek.
The crest elevation of the auxiliary spillway is set at the lake elevation corresponding to the 100-year flood level plus wave runup at the spillway due to a 40 mph overland I wind.
Design (SPF) discharge = 1,960 cfs Discharge due to SDF (PMF) = 15,530 cfs Crest elevation = 1,990.5 feet SL-3831
SARGENT A LUNDY W E N GIN E E R5 cnica 04-03-81 The following crosion protection provisions are provided for the auxiliary spillway so I that the dam embankment will not be endangered by PMF discharge:
- a. Tbc crest of the spillway and the downstream 3:1 slope are formed with a 1-foot thick concrete apron. The apron on the 3:1 slope continues down to the solid Toronto limestone rock at an approximate elevation of 1971 feet.
- b. The 10:1 slopes of both sides of the spillway along the axis of the dam also have a 1-foot thick concrete apron.
- c. On either side of the spillway, the concrete apron, with a 3:1 slope, continues through the transition of the dam and spillway channel up to the Toronto limestone rock at an elevation of about 1971 feet, as shown in Exhibit 4-1.
- d. The sloping concrete apron on the sides of the spillway channel continues downstream such that any erosion during PMF is checked about 100 feet I from the toe of the main dam, and thus does not endanger the dam embankment.
I e. The discharge channel and the sides of the channel are backfilled to the final grade with the excavated local material.
4.1.3 Reference Drawings Sargent & Lundy Drawings:
S-57 Service and Auxiliary Spillways - Plan 5 S-58 Auxiliary Spillway - Detail 1 S-61 Auxiliary Spillway - Plan, Profile, and Section 4.2 Structural Design 4.2.1 Design Criteria The auxiliary spillway, in conjunction with the service spillway, is provided to discharge the PMF with an antecedent SPF. A 1-foot thick concrete apron is provided to prevent possible dam erosion at the spillway due to lake discharge.
sir 3831
I SARGENT & LUNDY E NGIN E E RS CillCACO 04-03-81 I Becatse the apron provides no other structural function, a minimum area of reinforcing (No. 6 by No. 6/6-inch welded wire fabric) is provided as temperature and I shrinkage reinforcing.
4.2.2 Reference Drawings Sargent & Lundy Drawings:
S-6 Main Dam Sta. 0+00 to Sta. 60+65 S-57 Service and Auxiliary Spillways - Plan S-58 Auxiliary Spillway - Detail 1 S-61 Auxiliary Spillway - Plaa, Profile, and Section SARGENT & LUNDY Prepared by: 04-03-81 v
S. N. Kazmi Supervising Design Engineer Structural Design & Drafting Division il 04-03-81 G. V. Komanduri Supervisor, Water Resources Section Water Resources &
Site Development Division b- 04-03-81 R. D. Nelson Supervisor, Engineering Section i
Gectechnical Division cLY - Reviewed by: N J. M .' KutiTF-04-03-81
/
3 sNIIQ', ),,
Structural Project Engineer Structural Project Engineering I ! ,_.,I 7219 y Division w 15 f s Approved by: b
- 04-03-81
. k.;' Ch-AL[ SS}' N M. MeLaughliQ
, Manager,
! Structural Department 1
SL-3831 I
SARGENT& LUN DY W E N GIN E E R5 c m c ac.o 04-03-81 I
5.0 REFERENCES
- 1. Kansas Water Resources Board, "A Program of Fluvial Sediment Investigations in Kansas," Bulletin 6, July 1961.
I 2. Kansas Water Resources Board, " Sediment Yields from Small Drainage Areas in Kansas," Bulletin 16,1971.
- 3. " Proposed Guidelines for Safety Inspection of Dams," Department of the Army, Corps of Engineers, National Dam Safety Program, Federal Register, Vol. 39, No.168, August 28,1974.
- 4. Lambe, T. W. and Whitman, R. V., Soil Mechanics, John Wiley & Sons, Inc.,
New York,1969, Chapter 19.
- 5. U.S. Navy Design Manual, " Soil Mechanics, Foundations, and Earth Structures,"
NAVFAC DM-7, March 1971.
- 6. Sherard, J. L. et al, Earth and Earth-Rock Dam Engineering Problems of Design &
Construction, John Wiley & Sons, Inc., New York,1966.
- 7. U.S. Army Corps of Engineers, " Stability of Riprap and Discharge Character-istics, Overflow Embankments, Arkansas River, Arkansas," Publication No. 2-650, June 1964.
[ 8. Olivier, H., "Through and Overflow Rockfill Dams - New Design Techniques,"
l Proceedings of the Institute of Civil Engineers, Paper No. 7012, Vol. 36, March 1967.
- 9. U.S. Bureau of Reclamation, Department of the Interior, Design of Small Dams, Second Edition,1973.
- 10. Seed, H. B., Lee, K. L., and Idriss, I. M., "An Analysis of the Sheffield Dam Failure," Journal of the Soil Mechanics and Foundations Division, ASCE, Vol. 95, No. SMS, November 1969.
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SARGENT& LUN DY E N GIN E E R5 cmcaco 04-03-81
- 11. Seed, H. B., Lee, K. L., Idriss, I. M., and Makdisi, F., " Analysis of the Slides in the San Fernando Dams During the Earthquake of February 9,1971," Report EERC 73-2, Earthquake Engineering Research Center, University of California, Berkeley, March 1973.
- 12. Idriss, I. M., Lysmer, J., liwang, R., and Seed, H. B., "A Computer Program for Evaluating the Seismic Response of Soil Structures by Variable Damping Finite l
Element Procedures," Report EERC 73-16, Earthquake Engineering Research Center, University of California, Berkeley, July 1973.
- 13. Lee, K. L. and Chan, K., " Number of Equivalent Significant Cycles in Strong Motion Earthquakes," Proceedings of the International Conference on Micro-zonation for Safer Construction, Research and Application, Vol. II, November 1972.
- 14. Seed, H. B. and Peacock, W. H. , " Test Procedures for Measuring Soil Liquefaction Characteristics," Journal of the Soil Mechanics and Foundations Division, ASCE, Vol. 97, No. SM8, 8330, pp.1099-1119, August 1971.
- 15. U.S. Army Corps of Engineers Manual, " Engineering and Design - Hydraulic Design of Spillways," EM-1110-2-1603, Headquarters, Dept. of the Army, Office of the Chief Engineer, March 1965.
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T A BL ES I
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SARGENT O LUNDY ENG1NEERS TABLE 2-1 cmeaoo SL-3831 04-03-81 REQUIRED SAFETY FACTORS - SLOPE STABILITY ANALYSIS, CATEGORY l STRUCTURES Minlinum Safety Factors Minimum Safety Suggested in " Proposed Factor Required Guidelines for Safety at Inspection of Dams" Conditior. Wolf Creek Site (Reference 3)
- 1. End of Construction 1.4 No recommendation provided
- 2. Steady State Scepage (lleservoir at elevation 1970 feet) 1.5 1.5
- 3. Itapid Drawdown (elevation 1987 feet to elevation 1950 feet) 1.2 1.2
- 4. Earthquake (SSE) for Conditions 1 and 2 (Pseudostatic) 1.2 1.0
- 5. Earthquake (SSE) for Conditions 1 and 2 (Finite Element) 1.1 1.0 l
i l
SARGENT & LUNDV ENG1NEERS TABLE 2-2 cnicaoo SL-3831
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CHARACTERISTICS OF ONSITE AGGREGATE SOURCES i
i Toronto Plattsmouth Specific-Saturated 2.45-2.51 2.56-2.66 Gravity Specific 2.33-2.41 2.48-2.59 Gravity-Dry Los Angeles 31.4-38.2% 26.5-35.8%
Abrasion Test Absorption 2.46-5.26t 1.2-3.2%
Soundness .91 .96% .92 96%
Loss Ratio Note: Soundness Loss Ratio determined according to Kansas State
!!ighway Commission procedures.
Ref: (1) Stallard, A . II . , 1966, Materials Inventory of Coffey County, Kansas: prepared by the State liighway Commission of Kansas in cooperation with the U.S. Department of Commerce, Dureau of Public Roads.
I
M M M M M M SOll PARAMETERS FOR STABILITY ANALYSIS UHS DAM End of Construction Steady State & Rapid Drawdown
- Soil Cohesion Friction Angle Dentity Cohesion Friction Angle Density (psf) (Degree) (pcf) (psf) (Degree) (pcf)
{
i IEmbankment Soil 450 0 110 265 20 118 i
- Shale 1000 0 150 100 0 150 ,
I >
n :D Z$
oParameters are different from true values. However, they do not affect the solution. 90 2 Rz d !
E n IP Om
- C
- Z 0
l
@f5 eas
&$m Y
W a
M REPRESENTATIVE SOIL PROPERTIES FOR USE IN DESIGN ANALYSIS UES Category I Dam Parameter (Femolded Soils) In Situ Soils DENSITY (pcf)
Dry Density 90.00 97.50 Wet Density (as built or natural) 110.50 120.50 Wet Density (saturated) 118.00 124.00 Submerged Density 55.50 61.50
-7 PER?1EABILITY (cm/sec) 1.0 x 10 1.0 x 10 ~I m
COliPRESSIBILITY INDICES Preconsolidaticn Pressure (psf) 1500.00
,$0 3500.00 2 Cc/(1 + e) 0.13 0.16 go$
Cc/(1 + e) 0.01 0.04 52 H e 0.90 0.73 EmP on UNDRAINED SEEAR STRENGTF (psf) $ph At Natural or biolding floisture Contert 500 0 15 strain 1000 2 57 strein 0 At Saturated !!oisture Content 450 2 17 strain 7500 I 5< strain EFFECTIVE STRENGTH PARAF.ETERS Cohesion (psf) 266 peak 400 2 ST strain Friction Angle (degrees, 20 peak 20 a 5F strain
?h5 eas
&hm Y
Ww m
E SARGENT O LUNDY EN0lNCCRS TABLE 2-3.3 C HIC AGO $[.}8}j EFFECTIVE STRESS PARAMETERS MODIFIED MOHR DIAGRAM Test C1' + CO ' ' ~ G' Depth Pit (Pt.) 6 O u GT' =k-u 6' =k-u 2 2 TP-1 l'-3' O.84 0.30 0.22 0.62 0.08 0.350 0.270 1.17 0.60 0.45 0.72 0.15 0.435 0.285 1.60 0.90 0.65 0.95 0.25 0.600 0.350 0.78 0.30 0.20 0.58 0.10 0.340 0.240 1.62 0.90 0.58 1.04 0.32 0.680 0.360 TP-2 l'-4' O.88 0.30 0.13 0.75 0.17 0.460 0.290 1.29 0.60 0.30 0.99 0.30 0.645 0.345 1.77 0.90 0.50 1.27 0.40 0.835 0.415 0.89 0.30 0.14 0.75 0.16 0.455 0.295 1.72 0.90 0.50 1.22 0.40 0.810 0.410 TP-3 l'-3' O.78 0.30 0.22 0.56 0.08 0.320 0.240 1.64 0.90 0.54 1.10 0.36 0.730 0.370 1.20 0.60 0.40 0.80 0.20 0.500 0.300 TP-9 5' l.94 0.72 -0.17 2.11 0.89 1.500 0.610 3.04 1.44 0.17 2.88 1.27 2.075 0.800 3'-6' l.96 0.61 -0.06 2.02 0.67 1.345 0.675
" 0.395 3'-7' l.44 0.65 0.12 1.32 0.53 0.925 TP-10 0.820 I "
2'-3' 2.72 1.57 2.75 1.08 0.61 0.58
-0.06 0.65 0.19 2.78 0.92 2.56 1.14
-0.04 0.39 1.960 0.440 1.475 0.480 1.085 G]and u are at peak deviator stress. All units are tsf.
- 1.0 -
I S'= sin-1 (tan 0)=20
- 19 co^s e = 0.133 tsfcf265 psf C'= x
. a m
a a 3 g 0. 5 - x o TP-1 O A TP-2 e TP-3 o TP-9 x TP-lO o = 0.125 i i 1 I
1 0.5 1.0 1.5 2.0 t
71+6' (tsf) 2
SARGENT & LUNDY ENO1NEERS TABLE 2-4 C HIC AGO SL-3831 04-03-81 RESULTS Of UNCONflNED COMPRESSION TESTS ON ROCK CORE SAMPLES Static Unconfined Modulus of Compressive Elasticity
- I Geologic Unit Depth Strength Times x 10-6 Poisson's E91.i_D9_il n d L i t ho l ogy__j f e e tL__Jp s iL______Jp sL___ P a t i o *_
P-11 lieumader Sh. 25.3 69 0.00182 0.37 B-4 lieumader Sh. 22.9 300 0.0343 -
I P-4 !!eumader Sh. 36.3 56 0.00104 0.42 P-9 ileumader Sh. 35.8 131 0.00553 0.40 HS-28 Plattsmouth Ls. 33.0 6,690 9.26 0.27 HS-29 Plattsmouth Ls. 34.2 11,420 8.59 0.22 HS-28 Plattsmouth Ls. 42.8 5,380 4.78 0.29 HS-28 Heebner Sh. 45.5 1,110 0.110 0.30 HS-28 Leavenworth Ls. 47.9 10,910 8.43 0.24 lis-29 Leavenworth Ls. 49.0 11,710 9.89 0.16 HS-29 Snyderville Sh. 52.6 175 0.0127 0.32 HS-28 Snyderville Sh. 56.3 151 0.0140 0.31 HS-29 Toronto Ls. 61.8 2,910 0.852 0.30 HS-28 Toronto Ls. 63.8 6,760 3.85 0.24
- At 40 percent peak unconfined compressive strength.
I J
SARGENT & LUNDY ENG1NEERS TABLE 2-5 CHICAGO $[.3831 04-03-81 RESULTS OF SLOPE STABILITY ANALYSES FOR UHS DAM Factor of Safety Number Condition Static Pseudostatic Minimum Required
- 1. End of Construction 2.45 1.48 1.4
- 2. Rapid Drawdown from I clevation 1987 feet to elevation 1950 feet 2.18 -
1.2
- 3. Steady State Scepage, reservoir at elevation 1970 feet 2.50 1.57 1.5
- 4. Fully Submerged Condition, reservoir at elevation 1970 feet 4.67 2.09 1.5 I
SARGENTreLUNDY cyain,cas TABLE 2-6 cmc, "
SL-3831 04-03-81 SOIL PARAMETERS FOR STATIC STRESS ANALYSIS OF SUBMERGED UHS DAM Total Weight (pcf) = 118.0 Submerged weight (pcf) = 55.6 Effective Cohesion, c' (psf)= 265.0 Effective Angle of Internal Friction, t' (deg) = 20.0 Poisson's Ratio, p = 0.4 Modulus of Elasticity, = 50,000 E (psf)
I I
l
- I
(
=
SARGENT & LUNDY E N GIN E E R S TABLE 2-7 CHICAGO ${.}g 04-03-81 INITIAL STRESS AND FAILURE CONDITIONS Initial Stress Condition Failure Condition No. Sample K c "3c 0 .- U 3 Cyclic Axial Load 1 (tsf) (tsf) (tsf) Ado (tsf) { For 5 Cycles i 1 TP-3 1.25 0.6 0.75 0.6 0.780 2 TP-3 1.25 0.9 1.125 0.9 0.850 3 TP-13 1.75 0.2 0.35 0.2 0.615 4 TP-13 1.75 0.6 1.05 0.6 0.830 I l l 0 l l
SARGENT & LUNDY EN GIN EERs TABLE 2-8 CHICAGO SL-3831 04-03-81 CYCLlr ,R STRENGTH AND NORMAL STRESS FROM STRtsS-CONTROLLED DYNAMIC TRIAXlAL TEST Cyclic Shear Normal Stress o Strength o a =Tre No. Sample K c (tsS) I f (tsf) (tsE) tffc V 1 TP-3 1.25 0.6 0.360 0.650 0.108 2 TP-3 1.25 0.9 0.400 0.975 0.108 3 TP-13 1.75 0.2 0.293 0.250 0.288 4 TP-13 1.75 0.6 0.387 0.742 0.288
SARGENTQLUNDY E N G1N E E R S TABLE 2-9 c"'c^* st.3831 04-03-81 COMPUTED FACTOR OF SAFETY FOR THE FINITE ELEMENT MODEL OF UHS DAM Initial Vertical Initial Cyclic Induced Normal Shear Shear Shear T Element Stress Stress o Strength Stress F.S=T f No. co'(psf) To (psf ) co' Tf(psf) , Td (psf)_ T d 2 82.61 21.52 0.261 125.0 40.6 3.08 3 187.01 38.75 0.207 205.0 72.7 ' 82 4 293.85 57.85 0.197 295.0 117.4 .51 I 5 6 7 403.42 517.00 622.00 76.58 83.34 90.62 0.190 0.161 0.146 370.0 430.0 480.0 163.5 211.3 255.2 2.26 2.04 1.88 8 730.26 90.11 0.123 525.0 293.4 1.79 9 830.84 79.52 0.096 535.0 325.2 1.65 10 912.40 56.27 0.062 560.0 350.1 1.60 11 955.98 29.33 0.031 565.0 364.9 1.55 12 969.67 8.43 0.009 560.0 369.8 1.51 18 622.32 90.62 0.146 480.0 245.5 1.96 26 89.73 25.60 0.285 150.0 66.5 2.26 27 184.50 40.67 0.220 230.0 104.7 2.20 28 289.85 48.74 0.168 295.0 155.5 1.90 29 396.79 63.03 0.159 350.0 202.9 1.72 30 509.84 60.48 0.119 400.0 241.8 1.65 31 615.58 56.85 0.092 440.0 273.0 1.61 32 701.52 40.47 0.058 480.0 296.9 1.62 33 747.62 20.29 0.027 490.0 311.9 1.57 34 761.39 5.63 0.007 490.0 317.0 1.55 39 509.84 60.48 0.119 400.0 237.0 1.69 46 89.19 22.43 0.252 135.0 99.9 1.35 l 47 180.15 32.95 0.183 205.0 135.6 1.51 l 48 288.32 34.86 0.121 275.0 176.6 1.56 49 394.67 43.90 0.111 330.0 205.9 1.60 50 489.18 28.16 0.058 370.0 226.1 1.64 I 51 52 56 535.12 547.84 394.64 12.39 3.28 43.90 0.023 0.006 390.0 400.0 239.7 244.5 1.63 1.64 l 0.111 330.0 202.7 1.63
- 63 182.06 25.86 0.142 200.0 126.0 1.59 64 275.60 18.12 0.066 245.0 142.5 1,72
, 65 318.72 9.28 0.029 265.0 153.0 1.73 l 66 330.77 1.77 0.005 270.0 158.1 1.71 69 275.60 18.12 0.066 250.0 142.1 1,76 74 83.40 10.11 0.121 102.0 64.0 1.59 l
E M llNDRAINED STATIC STRENGTH AFTER DYNAMICALLY LOADING THE SAMPLE (c*) tn m 3 2 O Cyclic Axial Undrained c* no y 03c No. of Cycles Load c* c (static) c (sta tic) I -4 Sample Test # (psf) M N Adp (psf) _ (psf) (psf) (%) {2 Om g TU-13 600 p 1 1.0 11 400 390 580 67.2 2c
*Z TP-13 2 400 1.75 11 400 519 580 89.5 0 TP-13 3 1200 1.75 11 940 800 1030 78 hh5 eas to $ m Y
8
M M M M M l SOIL PARAMETERS FOR STABILITY ANAL YSIS ESWS PUMPHOUSE CHANNEL AND UHS SLOPES a8 End of Construction Steady State & Rapid Drawdown o *g Soil ' Soil Density Friction Angle Cohesion Friction Angle Cohesion hz" Number Description (pcf) (degree) (psf) (degree) (psf) g" 3C 1 Water 62.4 0 0 0 0 *z 2 Residual 124 10 585 20 400 3 Rock 160 35 5006 35 5000
??$
eas SOC s
i m l RESULTS OF SLOPE STABILITY ANALYSIS FOR UHS EXCAVATED SLOPE USING WEDGE ANALYSIS m Computed Minimum Required Minimum "b Z m Condition Factor of Safety Factor of Safety SS$ l, nz
- 1. En P End of* Construction 7.8 1.4 "r 3C Steady State
- 2. 5.3 1.5 *$-<
- 3. Steady State plus SSE (0.12 g) ,
3.5 1.2 hh$ eas
&Om Y
d
M M RESULTS OF SLOPE STABILITY ANALYSIS FOR ESWS INTAKE CHANNEL Computed M'nOsum i Factor of Safety Required Minimum Conditions Slope 3:1 Slope 5:1 Factor of Safety
- 1. Steady State 7.13 3.37 1.5 g S
- 2. Steady State plus SSE (0.12 g) 3.37 1.86 1.2 go$
5z d
!; m sp m
- 3. End of Construction F.S. with SSE of 0.12 g is 3.88 (higher than $ $r-1.5). Therefore, no o analysis is performed. 3.14 1.5
- P
- 4. End of Construction plus SSE (0.12 g) 3.88 1.74 1.2
- 5. Rapid Drawdown 5.69 ii.;: no effect on stability 1.2 l
l I l ( ??$
- 8. w$ rn E
a h) a w
SARGENTO LUNDY E Ed G l Gd E E R S TABLE 2-14 CHICAGO , 04-03-81 I RESULTS OF CONSOLIDATION TESTS ON UNDISTURBED AND RECOMPACTED SOIL SAMPLES Preconsolidt. tion Compressibility Swelling-I Location Depth (feet) Pressure (psf) Indexl (in/in) Index (in/in) Soil Type 3 4 CL TP-1 1.0-3.0 1600 0.124 - llSA-1 3.0 3600 0.134 0.050 CL llS-2 6.0 - 0.170 0.034 CL llS-5 2.0 4600 0.150 - CII I flS-16 4.0 8400 0.180 0.022 Cli IIS-17 4.5 8600 0.112 0.040 CL
- 1. Compressibility Index is defined as Cc/(1 + e).
- 2. Swelling Index is defined as Cs/(1 + e).
- 3. Sample compacted to 95% of maximum dry density and 4% wet of optimum moisture content as determined by ASTM D-698-70.
- 4. Apparent preconsolidation pressure.
i SARGENT & LUNDY ENGINEERS TABLE 2-15 C HIC AGO R-3W SHEET 1 OF 2 04-03-81 I RESULTS OF FIEl.D PERMEABILITY TESTS ULTIMATE HEAT SINK A. PRESSURE WATER-LOSS TESTING Average Permeability Number of I Member Permeability
- ___Jem/ s ec)
Range ____ (cm/ sect ___ Number of_Tgg__No Ia]Sgs_JO_1,_ Heumader 3.0 x 10-a 0 - 6.0 x 10-* 8 6 Plattsmouth Ls. 4.0 x 10-* O- 1.4 x 10-5 26 15 Heebner 9.0 x 10-6 0- 2.9 x 10-5 29 16 Leavenworth Ls. 7.0 x 10-6 0 - 3.6 x 10-5 29 13 Snyderville 9.0 x 10-6 0 - 4.8 x 10-5 36 17 Toronto Ls. 2.0 x 10-5 0- 1.0 x 10-* 22 6 I 10-s em/sec assumed when computing averages. I *O = No take recorded. I
SARGENT & LUNDY ENGINEERS TABLE 2-15 c"'c " SL-3831 SHEET 2 CIF 2 04-03-81 B. FALLING HEAD PERMEAMETER Slotted Boring Interval Members Permeability Number Piezometer (feet) MgDiggred _____Jcm/secl____ HS-1 A 2.5 - 19.5 Soil-Plattsmouth Ls. 6.1 x 10-* Snyderville-B 30.0 - 37.0 Toronto Ls. 2.7 x 10-* HS-3 A(test 2) 3.0 - 17.9 Plattsmouth Ls. 8.5 x 10-* Toronto Ls. HS-5 A 4.8 - 9.8 Plattsmouth Ls. 9.0 x 10-* Snyderville-B 23.5 - 30.3 Toronto Ls. 5.4 x 10-7 HS-8 A 4.8 - 9.8 Plattsmouth Ls. 4.4 x 10-* Snyderville-B 31.0 - 39.5 Toronto Ls. 2.2 x 10-6 HS-20 A(test 2) 2.0 - 18.0 Soil-Plattsmouth Ls. 7.6 x 10-* B 35.0 - 43.0 Toronto Ls. nil i I HSA-1 A B 3.0 - 11.5 15.0 - 22.2 Soil Toronto Ls. nil 1.7 x 10-5 C. CONSTANT HEAD PEREAMETER TESTING Slotted Boring Interval Permeability Number __JfeetL__ __Megber ___ (cm/sec) __ l HS-SP-2 2.5 - 7.0 Soil 1.2 x 10-8 HS-SP-4 2.0 - 11.5 Soil 1.0 x 10-* HS-SP-5 2.0 - 7.0 Soil 4.7 x 10-6 l HS-SP-6 2.0 - 4.4 Soil 6.5 x 10-* 1 1 I l l
M M M M M RESULTS OF LABORATORY FALLING HEAD PERMEABILITY TESTS ON UNDISTURBED AND RECOMPACTED SOIL SAMPLES Field or Initial Saturated Moisture Moisture Dry m Depth Head Permeability Content Content Density Soil L9981190 (1221L___llef tL_. _ _lsW_22CL_____lERIE20tL___1 percent L____ (pe f L____Iype ]$mo HS-1 2.5 17 8.8x10-8 26.8 27.5 97.9 CL $S$ nz HS-3* 1.0 30 1.1x10-' 25.0 - - CL $" E HS-3* 1.0 30 3C 2.7x10-8 25.0 27.5 92.1 CL *2 O HS-6* 10.0 20.6 5.6x10-8 15.2 - - CL 4 HS-6* 10.0 20.6 No Flow in 15.2 16.1 119.9 CL t 2 Days 1 TP-1* 1.0-3.0 30.3 2.2x10-e 17,7 inu.1 CL TP-1* 1.0-3.0 30.3 4.1x10-e 17.7 105.3 CL TP-3* 1.e-5.0 21.7 5.6x10-* 19.6 91.9 CH TP-3* 2.0-5.0 21.7 7.5x10-8 19.6 90.5 CH hh$ eas
&dm a e,a
SARGENT O LUNDY E N G1N E E R S TABLE 2-17 c"'C^ SL-3831 04-03-81 MONITORING SCHEDULE FOR UHS AND UHS DAM Vertical Horizontal Sediment UHS Phase Movement Movement Pads Profile (Note 1) (Note 2) (Note 3) (Note 4)
- 1. During filling of UHS to 1969.5 Monthly Initial Spot visual inspection -
II. UllS at 1969.5 At occurrence At occurrence Visual inspection - III. Filling of area downstream of UHS dam to 1969.5 Monthly At start of filling down-stream area - - IV. UllS at 1970 At occurrence At occurrence - Initial survey after filling when level is at approxi-V. UHS filling to 1975 Monthly At occurrence , - VI. UHS at 1975 At occurrence At occurrence Visual inspection - VII. Water level
~
!I l
>1975 Monthly Note 5 Visual inspectic n yearly When required, based on I visual inspec-tion of sedi-l ment pads Vill. Drawdown below l 1975 At occurrence At occurrence Visual inspection 1- at occurrence -
1 l Notes 1. Monthly until submerged. If no movement is noted, then yearly thereafter. i
- 2. Sight along horizontal movement hubs and measure offsets.
- 3. Visualinspection to measure sedimentation thickness 6 months and 12 months after filling of UHS. Diver can measure accumulated thickness of sediment on i pads.
!I 4. Initial survey after filling must be taken by the method planned to be used in i the future and to make a comparison to the survey data taken before filling. l
- 5. No horizontal measurements will be required of UHS dam after it is submerged.
Wt'
SARGENT & LUNDY E N G1N E E R s TABLE 3-1 c"'C^ SL-3831 SHEET 1 OF 2 04-03-81 LOADING AND SYMBOL DESCRIPTIONS 3 TV g - Weight of Water = 0.0624 K/ft 3 TV - Weight of Concrete = 0.15 K/ft C TV - Weight of 1120 Truck = 40 K il2O W3 - Weight of Soil 3 Granular = 0.12 K/ft 3 Cohesive = 0.115 K/ft TV SC
- Vertical Surcharge due to 1I20 Truck = 0.5 K/ft P g3 - Lateral Active Soil Pressure 2
Granular = 0.04 K/ft /ft 2 Cohesive = 0.08 K/ft /ft P pg - Passive Soil Pressure 2 Granular = 0.24 K/ft /ft Cohesive = 0.23 K/ft /ft P gg - Lateral Surcharge Pressure 2 Granular = 0.167 K/ft 2 Cohesive = 0.335 K/ft P gg - Ilydrostatic Pressure = 0.0624 K/ft2 /ft K A
- Coefficient of Earth Pressure Granular = 0.33
! Cohesive = 0.67
- Coefficient of Passive Pressure = 2.0 lI K" I.F. - Impact Factor = 0.3 (maximum)
II - IIcight of Cover of Fill Associated with Component under Investigation t - Thickness of Component under Investigation L - Length of Span of Component under Investigation l C.G.(A) - Moment Arm to Center of Gravity of Component "A" A1 0 - Ultimate Design Moment
I SARGENT & LUNDY ENG1NEERS TABLE 3-1 I CHICAGO SL-3831 SHEET 2 OF 2 04-03-81 51 C
- Ultimate Capacity Moment C.M. - Capacity Margin = MC I U f r Factored Loads c - Eccentricity p - Coefficient of Friction d - Effective Depth r
l t 1 I I l l l l i 1 l l
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t EXHIBIT 1-7 SL-3831 4 04-0.L81 1,.I1 1 99 d%m esat sham.aswis hs # - K w *' ~ 'l
~
-,,, ev, ,e. .r l ' Opt
[ El, _1 W 2[** ('aa* 8 l*[ WMT l .* 9 - .ee woes,O_. fine use.mo egy pay . T to 1 l#** M *ary_g_LM__q 8b,"I<^ ~~ hjg W- (Y '.
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- q. .
SARGENT & LUNDY ENGINEERC EXHIBIT 2-8 CHICAGO SL-3831 04-03-81 SAFE SHUTDOWN EARTHQUAKE - HORIZONTAL GROUND SPEClRA 8 _',g v : y x :x ::,,7-v, m . , , ,
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VERTICAL GROUND SPECTRA ARE BASED ON~ AO,F
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i l ARTIFICIAL ACCELEROGRAM FOR HORIZONTAL GROUND MOTION NORMALIZED ACCLLERAT10N
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aus em mum , 5 ! ARTIFICIAL ACCELEROGRAM FOR VERTICAL GROUND MOTION NORMALIZED l ACCELERATION l
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3 SARGENT& LUNDY E '" '"""* EXHIBIT 2-12 CH1CoGO SL-3831 I SHEET 1 OF 3 04-03-81 HORIZONTAL DESIGN RESPONSE SPECTRA FOR 0.12 g HORIZONTAL GROUND ACCELERATION (10% DAMPING) I 2.0 10 0.5 P<Rie0. sec. 0.2 01 0.05 0.02 I 20., x N- ^ s/N/ i 10.0 S WN
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l O.05 0.5 1.0 2.0 5.0 10.0 20.0 50.0 FREQUENCY. CPS
SARGENT & LUNDY EXHIBIT 2-12 cme" SL-3831 I SHEET 2 OF 3 04-03-81 (7% DAMPING) I ~ ' 2.0 10 0.5 2 O.1 0.05 0.02 s N- s N-g 2 - p A p 10 0 '- '
~~~ _
^ \
5.0 AannclAL ACCELEROGRAM S 0 4< N / \ l s '- s b s s \/ p 3 1.0 0 l 05 ^ #
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!, A s ^ '
*^
O.1 0.05 ' # 05 1.0 2.0 5.0 10.0 20.0 50.0 FREQUENCY. CPS
SARGENT & LUNDY ceoeeccns EXHIBIT 2-12 cmcaco St_3831 ; SHEET 3 OF 3 I 04-03-81 (5'X, D AMPlNG) PERISO. SEC. 2.0 10 0.5 0.2 0 .1 0.05 0.02 20.0 i '
/^
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- I J
g SARGENT & LUNDY EMGlNEERS EXHIBIT 2-13 CHIC AGO SL-3831 SHEET 1 OF 3 04-03-81 VERTICAL DESIGN RESPONSE SPECTRA FOR 0.12 g HORIZO 4TAL GROUND ACCELERATION (10% DAMPlNG) I PERISO. SEC. I 2.0 1.0 0.5 0.2 01 0.05 0.02
' 20.0
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SARGENT & LUNDY EMGIMEERS EXHIBIT 2-13 csicco SL-3831 SHEET 2 OF 3 04-03-81 (7% DAMPING) PERIOD. SEC. 2.0 10 0.5 0.2 0.1 0.05 0.02 20.0 - N 10.0 I ^ 5.0 ,. -- - - Nh RESPONSE SPECTRUM OF THE x
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/
0.5 1.0 2.0 5.0 10.0 20.0 50.0 FREQUENCY, CPS I
SARGENT & LUNDY E M GIN E E R S EXHIBIT 2-13 cmc 4co SL-3831 I SHEET 3 OF 3 04-03-81 (5% DAMPING) 2.0 1.0 0.5 0.2 01 0.05 0.02 I (( [ 20.0 l 10.0 , [ /
\m \ \ \ \ \_ \
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o i I L yc i e, li l. r l. v-I [_ te r. - m 91.
- 96- . . _ . o. f. p 74 ,, 76 ,_76 ,,77 b "**
4 y[ s e, (
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' I __e "2/____I 70 __ _ 3L . . _ F2 . _64 rs 63 .
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t I L i . .. f i f
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l
EXHIBIT 2-14 SL-3831 04-03-81 w E vu
,g ~ , e1gg og
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se ir ss ;r. ro r, ri 35m ~. ., N,
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f } SUBMERGED UHS DAM, FINITE ELEMENT REPRESENTATION FOR DYNAMIC ANALYSIS h { I SARGENT$LUNDY ; i - J ENGINE E f3S '. g
SARGENT & LUNDY EMGIMEERS EXHIBIT 2-15 CHICAGO SL-3831 l 04-03-81 1 CYCLIC SHEAR STRENGTH FOR 5% STRAIN AND 5 CYCLES VERSUS NORMAL EFFECTIVE STRESS l 16000 - 1 l g 3fC l Ofc _E12000 w to ( i T
== O 288 G
z ",, - *a0.0.108 2 y "'
~~p,-== 0.0 "e
y 800 -
" , ,,, g-" , , , ~~,.,,,
x 3 ,s', ',#, g a ,-
? s' a s
a s s y 400 u
/ ,' y '
f
/
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//
j i i i e i s O 400 800 1200 1600 2000 2400 .I NORMAL STRESS ON FAILURE PLANE DURING CONSOLIDATION, 0-{c (psf) l l I I I
a Y__ ._ Normal Main Cooling Lako El 198?'- - - - _V l 1987 _ 20' 1974 - 4
'O 1966 - 3l u
w
. Comp 0Cted Cic >- c i o i I, j 1958 -
a > o
, , , A-____________.____________ _-
-90 -8'2 -7'4 -6'6 $8 $0 -42 -34
-26 -18 -lo -4 0k 800 -
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l j_ $"E 400 5 E t$ m 5z 3 l.. ' ' ' ' ' ' " > > s a
< 0 -82
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9 .- ea
- 10 0 -
9 z "N , 4 k e . . . , o i
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l T' s O.2 -
- s. 6 N.
O I j 01 - 4 a: s e e i e i i i i i __ t o i
-82 -74 -66 - 58 -50 -42 34 -26 -18 -10 -4 0 0 j
h W t' I
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) N I2 G rn e 400
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5 u8 , , -
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e s I a 1 1 i f i a i 0 f 66 -58 -50 -42 -34 -26 -18 -10 -4
-82 -74 V
t Distance Fr o m Axis Of i_ , 6
- - -n. , -
i EXHIBIT 2-i6 SL-3831 3 _, 04-03-81 ll ly ______________________w_
~ ~[O~-Ya 26* ~ -I4 42 5d--$8 $6- 7'4 82 9O cr.' - 1 . -- - - - . -- - _ _
to 10 18 26 34 42 50 58 66 74 82
- l - --+ t + t-- i + l ; ;
Y _ g _.. -_ .-..- t 3 ; g Cyclic Shear Strength For 5% Strain 8 5 Cycles, Ig .... i
,,,,_gn d uce d Shear S tre ss
.s,orF 5 Cycles , Zd
,,,,*- DYNAMIC ANALYSIS OF SOIL STABILITY
. , , ,, ALONG THE BASE OF UHSD
~* L
_ _g _ _ _ i - g. _ - - _ _g ._ . . _ 2 5O ~$8 66 l4 4 (ELEVATION 1952 FEET) (m b a n k m e n t -(Ft) SARGENT&LUNDY
. ~ . , ~ . . . . _
l.
M M M RESULTS OF STATIC TRIAXIAL TESTS AFTER CYCLIC STRESSING { l e PHASE 7, TEST NO. 1 Y PHASE 7. TEST No. 2 SPECIMEN SIZE AT START OF DYNAMIC TEST: SPECIMEN SIZE AT START OF OYNAMIC TEST: AREA (Ao) : 6. 6 5 SQ . I N . AREA (Ao) : 6. 59 SQ. IN. HE IGHT (Ho) : 6.48 IN. HE f CHT (Ho) . 6.47 IN. LATERAL CONFINING PRESSURE (as) . 600 PSF LATERAL CONSOLICATION PRESSURE e (c ) . 400 PSF 5200.. CONSOLICATION STRESS RAT fD (Kc) : N.A. CONSOLICATION STRESS RATIO (Kc) : 3.75 CYCLIC AXI AL STRESS (aa. ) : ? 400 PSF CYCllc AX1 AL STRESS (aa. ) ! 400 PSF TOTAL NUMBER CYCLES: 11 TOTAL NUMBER CYCLES: Il MEAN TOTAL AX1AL STRAIN AT END OF DYNAMIC TESTS 0.04% MEAN TOTAL AX1 AL STRAIN AT END OF OYNAMIC TEST: 0.01% l
~
W PHASE 7, TEST No. 3 SPECIMEN SIZE AT START OF DYNAMIC TEST: AREA (Ao) : 6. 53 SQ. IN. 2m 2400- LAT AL CONS lbAT b P 5SURE (ac) : 1200 PSF CONSOLIDATION STRESS RATIO (Kc) . l.75 m a CYCLIC AX1AL STRESS (a a,) t 9'*0 PSF y
., TCIAL NUMBER CYCLES: Il M 21 rn MEAN TOTAL AXtAL STRAIN AT END OF DYNAMIC TEST: 0.2%
2 mo b 2000- "O I IZy nz {
$M@
on p m- 2C mz y 1600 - o e < H m
$ 1200-W H
m O N 800 -
=
a 400-
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$$E O . i i i e i s 8 8 I
&2 E O 2 4 6 8 10 12 14 16 18 20 $
N AXI A L ST R AIN (PERCENT)
?
i i C f SOIL PAR AMETERS SOfL DESCRIPTION REMARKS N2 df(pcf) c(g,f ) g 1 COMPACTED CLAY 118.0 380.O O PARAMETERS ARE 2 3 HALE 150.O 1000.O O RUE V S BUT THEY DO NOT
" O AFFECT THE 3
l jLUTION
' 4 a
- REDUCTION IN STRENGTH FROM DYNAMIC AND STATIC TEST
- MINIMUM SAFETY FACTOR: 2.07 4.
t
$. g 20' , 1970.0 i
19703 4 E ! l
' 3 5 l SOIL- I .e . ,
i
* $ 1950 SOIL - 2 a
2 z 9 Ell 947.5 ] F ti C d ,,30) EL.19 37 5 e 4 4 f
s EXHIBIT [-18 SL-3831 1 04-03-81 2 J5 2 96 2.07
/
/
7
/
/
2.'37
\ 2.'78 / /
.b i 0IL-3 SQL-4 SLOPE STABILITY ANALYSIS FOR UHS DAM STATIC STRENGTH FOLLOWING CYCLIC LOADING SARGENT&LUNDY'
, ~o ,~ . . . m W
SARGENT & LUNDY ENG1NEERS EXHIBIT 2-19 c"'c^ i SL-3831
, 04-03-81 WEDGE ANALYSIS OF EXCAVATED SLOPES - ULTIMATE HEAT SINK 4 /o M AXIMUM NORMAL EL 1980.O' FINAL GRADE \ FORCE-~;
EL1970 0' HEAT SINKS - SE FO i DRi{NG_ FORCE ___ t r- ' FRICTIONAL s EL1960.O' RE SISTANCE ROCK SURFACE WEDGE 30' O 30' 60' SCALE IN FEET I I I I
j ,..- h-%_ i r- SOIL PARAMETERS I SOIL
; SOIL UNIT WT. COHESION FRICTION ANGLE DESCRIPTION (T PCF C PSF 0 CEGREE i
1 DUMMY LAYER 2 WATER 62.4 0 0 3 RESIDUAL 12 4 400 20 4 ROCK 15 0 5000 35 MINIMUM SAFETY FACTOR = 7.13 i L 7. r# e, 4 so t .4 - 1994.00 - 2 a rt y Q iba g gi
& 5 0
z ,974.00_ u- 0 E L.1970 , ' 55, z i 9 Q S0ll 3 3
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a a i EXHIBli 2-2g SL-3831 04-03-81 12.57 19.93 SOIL I 25 86 V su 3
\ S0ll 2 L EL.1965 1 SLOPE STABillTY ANALYSIS,3:1 SLOPE,
' ESWS INTAKE CHANNEL STEADY STATE CONDITION SARGENTbLUNDY 9
=m. _..
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! SOIL UNIT WT. COHESION FRICTION ANGLI DESCRIPTION (T PCF C PSF 6 DEGREE i DUMMY LAYER 2 WATER 62.4 O O 3 RESIDUAL 12 4 400 20 4 ROCK ISO 5000 35 MINIMUM SAFETY FACTOR = 3.37 4
1-r 2034.00 - E' - 2014.00 -
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EL.1970, ! 55' SOIL 3 i 1954.00 - , I~ l l I SOIL 4 i 1934.00 _
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i EXHIBIT 2-Il ~ SL-3831 J 04-03-81 3.39i i3.52 3 37 3.7 3'38
= a h 3
\ ~
SOIL 2 L E L.1965 SLOPE STABILITY ANALYSIS,3:1 SLOPE, ESWS INTAKE CHANNEL STEADY STATE WITH SSE SARGENT&LUN0Y
==a.~~...._
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W a i W e. SQL-3 1954 -
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EXHIBIT 2522 SL-3831 04-03-81 ) SOIL PAR A M ET E RS SOIL DESCRIPTION UNIT WT. COH E SIO N FRICTION ANGLE (T PCF C PSF 6 DEGREE I DUMMY LAYER 2 WAT E R 62.4 O O 3 RESIDUAL 12 4 585 10 4 ROCK 15 0 5000 35 MINIMUM SAFETY FACTOR = 3.88 3.90 4 03 388 4 26 34
/ \
/ SOIL - 1 l
\ SOIL- 2
-4 SLOPE STABILITY ANALYSIS,3:1 SLOPE, ESWS INTAKE CHANNEL END OF CONSTRUCTION WITH SSE SARGENT&LUNDY
' E NGIN E E f85 a I.
i
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t < t i . L l 8 SOIL PAR AMET ERS SOIL DENSITY COHESION FRICTION A DESCRIPTION { / PCF C PSF G DEGRE t. I DUMMY LAYER O O O 2 RESIDUAL I24 585 10 [ 3 ROCK 150 5000 35 il MINIMUM SAFETY FACTOR = 5.69 e E e 4 m a 3 i r-g_ _ 1995 - 2 s
, y 1986 _f O v 11
% 5 Z 1975 -
{
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1970 U _ z o SOIL 2 I U
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d 1955 - gg SOIL 3 k t 6 , a 9
EXHIBIT 2g23 SL-3831 . 04-03-81 i NGLE E ( OF CHANNEL 6 67 6.46 55.0 FT. ! 5.69 698 SOIL l g A%%\\ SLOPE STABILITY ANALYSIS,3:1 SLOPE, ESWS INTAKE CHANNEL RAPID DRAWDOWN CONDITION SARGENT&LUNDY
- 1. ~ o,~. . '. a I
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$ 2 015 -
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EXHIBIT 2-24 SL-3831 I 04-03-81 SOIL PARAMETERS Soil SOIL DENSITY COHESlON FRICTION ANGLE DESCRIPTION (T PCF C PSF 6 DEGREE I WATER 62 4 O O 2 RESIDUAL 12 4 400 20 3 ROCK 15 0 5000 35 MINIMUM SAFETY FACTOR = 3.37 Q, OF CHANNEL - 50.0 FT. l Y Y SOIL l 1
@XR\
SLOPE STABILITY ANALYSIS, S:1 SLOPE, ESWS INTAKE CHANNEL STEADY STATE CONDITION SARGENT&LUNDY r
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t j i b 9 i {. 2055 - n 6 1 87 189 t. 2035 ~~ t96 1 t6 - g 87 1 1 0 ( U O 2015 ~ 4 b W i . 1-o
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s EXHIBIT 2-2$ SL-3831 4 04-03-81 SOIL PARAMETERS SOIL DENSITY CO H E S IO N FRICTION ANGLE DESCR TION (T PC F C PSF 0 DEGREE l WATER 62.4 0 0 2 RESIDUAL 12 4 400 20 3 aOcx 15 0 5000 35
/. 86 MINIMUM SAFETY FACTOR N
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- 55.0 FT. l i SOIL I I WA\\ SLOPE STABILITY ANALYSIS,5:1 SLOPE, ESWS INTAKE CHANNEL STEADY STATE WITH SSE SARGENT&LUNDY)
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I EXHIBIT 2-26 SL-3831 1 04-03-81 SOIL PARAMETERS SOIL SOIL DENSIT Y COMESION FRICTION ANGLE DESCRIPTION
/ TPCF C PSF 0 DEGREE I WAT ER 62.4 0 0 2 RESIDUAL 124 585 10 3 ROCK ISO 5000 35 MINIMUM SAFETY FACTOR = 3.14
( OF CHANNEL - 55.0 FT. ! Y Y
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SOIL l 1 SLOPE STABillTY ANALYSIS,5:1 SLOPE, ESWS INTAKE CHANNEL END OF CONSTRUCTION
=-
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EXHIBlTl2-27 SL-3831 04-03-81J SOIL PARAMETERS SOIL SOIL DENSITY COHESION FRICTION ANGLE DESCRIPTION PCF C PSF 0 DEGREE I WATER 62.4 0 0 2 RESIDUAL 124 585 10 3 ROCK 150 5000 35 MINIMUM SAFETY FACTOR = l.74 ( OF CHANNEL - 55. 0 FT. l V Y h 3 SOIL I 1 I
/
t SLOPE STABILITY ANALYSIS,5:1 SLOPE, ESWS INTAKE CHANNEL END OF CONSTRUCTION WITH SSE SARGENT&LUNEY
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j i l i 80 i e
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SL-3831 04-03-81 I i I l DESCRIPTION OF COMPUTER PROGRAMS REFERENCED IN THE REPORT I I I I 1I I I I I I l I
SARGENT & LUNDY EMGINEERS CHICAGO BISHOP I BISHOP (Slope Stability Analysis) uses the Simplified Bishop - I Method to perform slope stability analysis. The factor of safety is defined in terms of moments about the center of the failure arc. The resultant of all forces on the sides of any slice is assumed to act horizontally. The pseudo-static approach is used to simulate the effect of an earthquake loading on the stability of slopes. The static equivalent earthquake force for each slice is applied horizontally through the center cf the base of that slice. Input to the program consists of slope geometry, soil character-istics, ground water level, centers and radii of trial circles, and the number of slices to be used in the analysis. Either SI or U.S. customary units can be used. Output from the program includes an echo print of the input data and the factor of safety for each trial circle. A plot of slope geometry and trial circles can also be produced. BISHOP was originally developed by J. E. Bowles of Bradley Univer-sity [1]. It was modified by Sargent & Lundy in 1975, and the POL (Problem Oriented Language) was added in 1976. The program is now maintained on UNIVAC 1100 series hardware operating under EXEC 8. A typical slope cross section was used for validation. The BISHOP results were compared with results from the ICES-SLOPE program [2]. 1
-- )
SARGENT & LUNDY ENG1NEERS CHICAGO I The slope geometry of the problem and the soil properties used in the slope stability analyses are shown in Figure 1. The resulting factors of safety for three loading conditions obtained from BISHOP and from ICES-SLOPE are shown in Table 1. The results correlate well. I i 2 i I __ _ - - - - . - .
1 ll l1 0 8 2 2 I 0 0 0 0 0 , H 2 L P I o 0 S f 3 N s 2 O p d I 5 0 0 0 0 nn S 6 0 0 0 aw E 2 0 0 0 k o H 1 1 1 n ed O i t w C S aa t r SD f t Y c 4 a yd T p . e m di I 8 0 0 0 2 H e ap S 1 5 5 5 6 o l ea N 1 1 1 1 d e b tR E t o S D 2 a r mP i t n L l o I 1 2 3 4 5 U i O t S f a od i e l p a oV l 0 S P _ 1 3 4 5 O
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I Table 1 Results of Problem Solved with BISHOP and ICES-SLOPE (BISHOP) p rc e Details Factor of Safety f) center radius BISHOP ICES-SLOPE x y R End of Construction 175 1125 75 1.67 1.67 Steady State Seepage with .lg 160 1145 95 1.98 2.02 Rapid Drawdown 175 1140 90 2.53 2.51 l I l 4 I
REFERENCE
- 1. J. E. Bowles, Analytical and Computer Methods in Foundation Engineering, McGraw-Hill Book Company, New York, 1974, pp. 465 and 467.
- 2. " ICES-SLOPE - Slope Stability Analysis System," McDonnell Douglas Automation Company, 1974.
I 5 Final I -
ISBILD ISBILD (Analysis of Stresses and Movements in Embankments) is a , finite element program developed to analyze static strains, stresses, and displacements in an embankment-foundation system using the latest finite element analysis techniques. The program takes into account the incremental loading to simulate the successive construction stages of an embankment. It employs nonlinear, hyperbolic and stress-dependent stress-strain behavior on the primary loading and stress-dependent stress-strain behavior on unloading and reloading. The program is an improved version of computer program EMBANK, developed by Kulhawy, Duncan and Seed in 1969 [1]. It was originally written by Ozawa and Duncan [2] of the University of California, Berkeley. Isoparametric elements with incompatible displacement modes [3] and a more accurate procedure for assign-ing the initial stresses are used in the ISBILD program. It also incorporates more efficient computational techniques including a new method of defining the boundary conditions and a new equation solver developed by Wilson in 1971 [4]. ISBILD can be used to calculate stresses and movements in a homogenous or zoned embankment on a rigid or stratified soil foundation due to gravitational or applied loads. It can also be used on a natural or cut slope by the gravity turn-on method provided in the program. The program works on two-dimensional problems assuming plane strain and isotropic conditions. 1
I ISBILD was acquired by S&L from the University of California at Berkeley and converted for use on the in-house UNIVAC 1100 series hardware. It was modified to include the capability to treat linear soil properties, to compute horizontal body forces and to calculate the movements in an embankment due to water loading. To validate Sargent & Lundy's UNIVAC 1100 version of the program, problems provided by the original authors were used as well as problems taken from published literature. Two of these problems are presented. In the first problem the measured movements of the Otter Brook Dam [5] are compared with the results of ISBILD. Figure 1 shows a comparison of the results for the horizontal displacement of the upstream face and for the displacement of the bridge pier. As shown the calculated results compare favorably with the field measurements. In the second problem a homogeneous dam resting on impermeable base with a full pool (Figure 2) is analyzed. Deflection distributions with depth are plotted for the centerline and the upstream and downstream faces (Figure 3). Solutions from ISBILD agree closely with those obtained by Carter, et al. [6]. 1 I 2 I
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(b) Horizontal Movement (positive downstream) l V'=o.4 Ce.r t c r , <f af
- ISBILD I Figure 3 Cc:..parison of water Load Effects on Movement (ISBII,0 vs. Reference 6) ;
I REFERENCES
- 1. F. H. Kulhawy, J. M. Duncan and H. B. Seed, " Finite Element Analysis of Stresses and Movements in Embankment During Construction," Geotechnical Engineering Research Report No. TE-69-4, Department of Civil Engineering, University of California, Berkeley, November 1969.
- 2. Y. Ozawa and J. M. Duncan, "ISBILD, A Computer Program for Analysis of Static Stresses and Movements in Embankments,"
Geotechnical Engineering Research Report No. TE-73-4, Department of Civil Engineering, University of California, Berkeley, December 1973.
- 3. O. C. Zienkiewics, B. M. Irons, J. Ergatoudis, S. Ahmad, and F. C. Scott, "Isoparametric and Associated Element Families for Two- and Three-Dimansional Analysis," Chapter 13 in " Finite Element Methods in Stress Analysis," Edited by I. Holand and K. Bell, Technical University of Norway, Tapir Press, Norway, Trondheim, 1969.
- 4. E. L. Wilson, " Solid SAP, A Static Analysis Program for Three-Dimensional Solid Structures," SESM Report 71-19, Structural Engineering Laboratory, University of California, Berkeley, September 1971.
- 5. K. A. Linell and H. F. Shea, " Strength and Deformation Characteristics of Various Glacial Tills in New England,"
Proceedings, Research Conference on Shear Strength of Cohesive Soils, ASCE, Boulder, Colorado, 1960, pp. 275-314.
- 6. J. P. Carter, H. G. Poulos, and J. R. Booker, "Effect of Seepage on Embankment Deformations Due to Water Loading,"
PB-242 450, National Technical Information Service, U. S. Department of Commerce, 1974. l 1 6 Final l l l
CARIENT O LUNDY ENGINEERO CMICAGO QUAD 4 I QUAD 4 is a finite element program which evaluates the seismic response of soil structures using a different damping ratio for each individual element. The base motion can be applied simul-taneously in two orthogonal directions. In addition, the procedure allows incorporation of both stiffness and damping values, that are strain-dependent, for each element. I I The program has been written for elements in plane-strain; triangular and quadrilateral elements can be used in represent-ing the continuum. The solution proceeds by assigning modulus and damping values to each element. Because these values are strain-dependent, an iteration procedure is adopted. Thus, at the outset, values of shear moduli and damping are estimated and the analysis is performed. Using the computed values of average strain developed in each element, new values of modulus and damping are determined from appropriate data relating these values to strain. Proceeding in this way, a solution is obtained incorporating modulus and damping values, for each element, which are compatible with the average strain developed. l l The program output includes the strain compatible soil prop-l l ! crties, response spectrum at specified nodes, probable maximum nodal accelerations, and probable maximum direct stresses, l shear stresses and shear strains for all elements. The shear stress time history can also be obtained for specified elements. i 1
CAR 2ENT Q LUNDY E NolN E E RO CHICAGO QUAD 4 was originally developed by I. M. Idriss, J. Lysmer, R. Hwang and H. B. Seed of the University of California, l Berkeley . It was modified and is now maintained by Sargent & Lundy. The program is currently processed at Sargent & Lundy on UNIVAC 1100 series hardware operating under EXEC 8. It has been operational since November 1973. To validate Sargent & Lundy's version of QUAD 4, a sample problem was taken from the original program documentation.1 A 100-ft. layer of dense sand shown in Fig. 1 has been analyzed. The properties of the sand were considered to be as follows: Total unit weight = 125 pcf (K2) max = 65 I Kg = 0.5 The parameter (K I relates the maximum shear modulus, G 2 max max' and effective mean pressure at any depth, y, below the surface as follows: G = c max 1000 (K 2 ) max m' ! (1 + 2K ) where o' = o' and K g
= coefficient of lateral pressure at rest and o' = effective vertical pressure at depth y.
Damping values and variations of modulus values with strain were based on data given in Reference (1). 2
CARCENT O LUNDY ENotNE3RO cricAeo The response of the sand layer was evaluated using the time history of accelerations recorded at Taft during the 1952 Kern County earthquake as base excitation. The ordinates of this time history were adjusted to provide a maximum acceleration of 0.15g. The sand layer has been represented by the finite element mesh shown in Fig. 1, which consists of 20 elements and 42 nodal points. To simulate a semi-infinite system, nodal points 1 through 40 have been fixed in the vertical direction and, there-fore, are only permitted to move in the horizontal direction. Nodal points 41 and 42 are fixed to the base. Comparison of results obtained from this validation run and the published results are shown in Figures 2-4. The values of damping and modulus, compatible with the strain level computed in each element, are presented in Fig. 2. The variations of maximum shear stresses and the maximum accelerations with depth are shown in Fig. 3. The acceleration spectrum for the computed surface motions is shown in Fig. 4. As illustrated by these figures, the comparisons are favorable. I I 3 i
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C Mt 0 0 b E P I R r fo)1 t G a e N I rc t n _M P M c e e r A p e D 1 S f e L " 0 nR _M . A o i . R O t s T a v C r E e4 P O l D S eA cU cQ A( C 3 0 2 8 4 O 0 6 O _M 2 1 I 0 O ' 4 o ZO~kaodOU< k&FdS t e r u g i F N M k
CARGENT O LUNDY ENGINEERS CHICAto I REFERENCES
- 1. Idriss, I. M., Lysmer, J., Hwang, R., and Seed, H. B.,
" QUAD 4 - A Computer Program for Evaluating the Seismic l
Response of Soil Structures by Variable Damping Finite I Element Procedures," Report No. EERC 73-16, Earthquake Engr. Research Center, University of California, Berkeley, California, July, 1973. l
- I l
l l l I i I 8 i l Final I .
1 I CARGENT O LUNDY E N GIN E E R S CHICAGO i RSG RSG (Response Spectrum Generator) generates dynamic response spectra (displacement, velocity and acceleration) for single-degree-of-freedom elastic systems with various dampings, subjected to a prescribed time dependent acceleration. The program may also be used to obtain a response spectrum consistent accelera-tion time history in which the response spectrum of the generated acceleration time history closely envelops the given spectrum. The dif ferential equation of motion is solved by neuraark's 8-method of numerical integration [1]. The program has the capability to apply a baseline correction in an earthquake acceleration time history as well as to obtain I and to plot the Fourier transform of the given acceleration time history. Options are available to obtain plots of the given acceleration time history, the generated response spectra alic their envelopen. In addition the respcmse c;>cetra can be combined using the probability method, square-root-of-the-sum-of-the-squares (SRSS) and absolute sum methods. An interpolation option to obtain an acceleration time history at equal intervals or at a smaller time interval is available. The program can also be used as a postprocessor for other programs with all its optiens and capabilities. Depending upon the option, the program output includes the response spectrum, the Fourier transform of a given acceleration time history, or the response spectrum consistent acceleration 1
~
l I CARGENTQ LUNDY ENGINEERO CHICAto time history. RSG was developed by Sargent & Lundy in 1969. Since 1972, 3 the program has been maintained on UNIVAC 1100 series hardware operating under EXEC 8. To illustrate the validity of the program, three sample problems are presented. In the first problem, response spectra of the El Centro north-south earthquake record (53-76 seconds duration) are generated using RSG for 0, 2, 5, 10 and 20 percent damping ratios. A time history plot of the earthquake record from RSG and as published in Reference [2] is shown in Figure 1. A comparison of response spectrum values obtained from RSG with the response spectrum values published in Reference [3] is shown in Table 1. Comparisons of response spectra plots at varying dampings are shown in Figure 2. As shown by the comparison, results obtained from RSG are accurate. The second validation example is a Fourier transform plot of a given 5 cycles /sec sine wave time history from RSG. The Fourier transform plot shown in Figure 3 shows a peak only at 5 cycles /sec. For the third validation problem, a spectrum-consistent time history was generated. A comparison of the desired response spectrum and the response spectrum of the compatible time history is shown in Figure 4. As seen from this figure, a good match is obtained. 2 {
, ' 'i tili ',
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I Table 1 Comparison of Response Spectra Values f rom RSG and Reference 3. Response Spectrum Values in (G) Units. No. Periods 0% Damping 2% Damping 5% Damping 10% Damping 20% Damping (sec) RSG REF(3) RSG REF(3) RSG REF( 3) RSG REF(3) RSG REF(3) 1 0.042 0.783 0.765 0.409 0.421 0.367 0.374 0.349 0.360 0.348 0.350 2 0.046 0.990 1.05 0.430 0.427 0.388 0.384 0.371 0.371 0.348 0.358 3 0.050 1.015 0.951 0.540 0.571 0.461 0.467 0.404 0.407 0.359 0.369 4 0.055 1.044 1.14 0.585 0.557 0.L28 0.416 0.403 0.406 0.365 0.382 5 0.070 0.963 0.803 0.471 0.493 0.439 0.446 0.424 0.424 0.383 0.398 6 0.080 1.333 1.32 0.708 0.712 0.582 0.579 0.482 0.488 0.390 0.407 7 0.085 1.070 1.07 0.682 0.668 0.591 0.591 0.486 0.490 0.395 0.408 8 0.100 1.755 2.07 0.815 0.805 0.567 0.567 0.480 0.480 0.409 0.419 9 0.130 1.990 1.99 1.040 1.03 0.773 0.772 0.535 0.541 0.429 0.445 I 10 11 12 0.150 0.180 0.200 1.977 1.309 1.609 1.92 1.49 1.58 0.847 0.887 0.916 0.837 0.890 0.914 0.578 0.726 0.650 0.579 0.727 0.644 0.497 0.587 0.531 0.504 0.597 0.542 0.435 0.452 0.444 0.454 0.474 0.463 13 0.220 2.415 2.50 0.731 0.728 0.683 0.667 0.580 0.576 0.442 0.471 14 0.260 1.538 1.60 1.177 1.15 0.903 0.902 0.639 0.653 0.443 0.469 15 0.280 1.309 1.11 0.878 0.882 0.755 0.746 0.567 0.578 0.430 0.463 16 0.320 1.820 1,82 1.067 1.07 0.699 0.703 0.519 0.527 0.402 0.432 I 17 18 19 0.360 0.380 0.400 1.212 1.717 1.964 1.21 1.72 1.99 0.877 0.978 0.824 0.877 0.972 0.827 0.657 0.673 0.614 0.655 0.678 0.615 0.504 0.493 0.473 0.511 0.503 0.481 0.379 0.390 0.409 0.408 0.403 0.418 20 0.440 1.591 1.58 0.969 0.966 0.728 0.731 0.553 0.558 0.463 0.478 21 0.480 1.405 1.41 0.996 0.996 0.794 0.797 0.644 0.651 0.514 0.537 22 0.500 1.179 1.17 1.018 1.02 0.830 0.836 0.691 0.699 0.533 0.559 23 0.550 1.988 1.99 1.266 1.26 0.910 0.917 0.745 0.759 0.545 0.588 I 24 25 26 0.600 0.700 0.800 1.252 1.846 1.089 1.25 1.84 1.08 0.970 0.898 0.671 0.971 0.900 0.670 0.854 0.619 0.547 0.859 0.622 0.549 0.706 0.534 0.436 0.722 0.546 0.444 0.516 0.408 0.307 0.570 0.459 0.347 I 27 28 29 0.900 1.000 1.200 1.176 0.829 0.818 1.17 0.83 0.818 0.754 0.676 0.440 0.755 0.677 0.441 0.536 0.515 0.330 0.539 0.518 0.331 0.385 0.350 0.236 0.393 0.359 0.241 0.267 0.231 0.173 0.289 0.249 0.186 30 1.400 0.420 0.420 0.237 0.237 0.181 0.181 0.170 0.173 0.130 0.138 I 31 32 1.600 1.800 0.327 0.490 0.327 0.503 0.242 0.230 0.243 0.230 0.194 0.178 0.195 0.179 0.159 0.146 0.162 0.149 0.124 0.122 0.136 0.136 33 2.000 0.353 0.353 0.226 0.226 0.178 0.178 0.148 0.152 0.120 0.135 I 4 {'.) .
1 QSG VALID.EXMPL.R/S ELCENTRS-40 T/H SARGENT 4 LUNDY oESiosER CsECnER 1 ENGINEERS ACTURL SPECTRA AT JOINUSLAB I 15 FEB 78 996CCS DAMP!NG 0.000 0.020 0.050 0.100 0.200 PAGE OF
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j El Figure 2 Comparison of Response Spectra Plots at Various Dampings from RSG and as Published in Reference 3. 1 5 C. S ese.v-Ah5/0378
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.9ine Wave Time Historv 6
3 FREQUENCY. CPS 50.0 20.0 10.0 5.0 2.0 1.0 0.5 50.0 I 20.0 10 0 5.0 e I { 2.0 E j Desired Spectrum e o 10 / Response Spectrum of
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I 0.2 01 I 0.05 0.02 0.05 0 .1 0.2 0.5 1.0 2.0 PERIGO, SEC. l SPECTRUM CONSISTENT TIME HIST. GENERATION- NSIG=2 Figure 4 Comparison of Desired Response Spectrum and Response Spectrum of Compatible Acceleration Time History (Damping = 0.02) from RCC 7
SARGENT O LUNDY E N GIN E E DS CHICAGO I REFERENCES
- 1. N. M. Newmark and E. Rosenblueth, Fundamentals of Earthquake I Engineering, Prentice-Hall, Inc., Englewood Cliffs, N. J.,
1971, p. 15.
- 2. Strong Motion Earthquake Accelerograms, Digitized and Plotted Data, Vol. II - Corrected Accelerograms and Integrated Ground Velocity and Displacement Curves, Part A - Accelerograms IIAOUl through IIA 20, California Institute of Technology Earthquake Engineering Research Laboratory, EERL 71-50, Pasadena, Califor-nia, September 1971.
- 3. A. G. Brady, et al, " Analysis of Strong Motion Earthquake Accelerograms, Volume III, Response Spectra, Part A, Accel-erograms IIA 001 through IIA 020," Prepared for the National Science Foundation, August 1972.
I I l I ( t l t I l l I l l l i 8 Final
SEEPAGE SEEPAGE (Two-Dimensional Steady-State Seepage Analysis Program) is a finite element program developed for analyzing various types of two-dimensional steady seepage flows through nonhomo-geneous anisotropic porous media, such as flow through an earth dike; flow into wells; and seepage losses through a bed of canals, lakes, etc. The program is capable of computing the pressure, potential function, stream function values, velocities in two directions on a vertical plane, and discharge values through vertical section lines in the flow domain. It can also determine the position of the fru. surface line and plot the flow net. Input for this program consists of the geometry of the flow domain, directional permeability coefficients, and available pressure heads on the boundaries. Output consists of nodal point pressures, potential values, stream function values, velocities and hydraulic gradients in two directions in every element, and discharge through specified sections. For seepage problems involving free surface, additional input is required, including the initial trial free surface, number of iterations for free surface, free surface correctior. tactor and error tolerance. SEEPAGE was originally developed by Robert L. Taylor of the University of California at Berkeley [1]. It has been extensively modified by Sargent & Lundy since 1972. It is now maintained at Sargent & Lundy on UNIVAC 1100 series hardware operating under I EXEC 8. I 1 I
1 To validate SEEPAGE, an axisymmetric flow problem and a plane flow problem are presented. An axisymmetric flow problem considering groundwater flowing into a well was analyzed by SEEPAGE and compared with hand calculations. The hand calculations are based on the well formula for steady radial flow in an unconfined aquifer as given in Reference 2. Figure 1 shows the finite element mesh configuration and permeability coefficients. The discharge obtained from SEEPAGE is 0.6791 cfs and that from the hand calculations is 0.6567 cfs. For the plane flow problem, Figure 2 shows a concrete dam resting on an isotropic soil having a permeability of 1.67 x 10 -5 (3), The results from SEEPAGE compare well with those from Reference 3 and with hand calculations based on the method of fragments (4] as shown in Table 1. 1 l !I 2 'I
I I I I of well ground
// 2 ,s u, </J Ng: ,,. s\ = //e %. ,,3 ss~
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M M M M L 6o' i
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i l I ! Table 1 Comparison of Results for Plane Flow Problem i
-SEEPAGE vs. Refs. 2& 3 i
l Method Discharge, ft 3/ min./ft Exit Gradient l SEEPAGE 6.63 x 10-3 0.45 l Lambe & Whitman 6.66 x 10-3 0.42 IIarr 7.27 x 10-3 0.48 i l I 5 I
REFERENCES
- 1. R. L. Taylor and C. B. Brown, " Darcy's Flow Solution with a Free Surface," Journal of the Hydraulics Division, ASCE, Vol. 93, No. HY2, March 1967, pp. 25-33.
- 2. V. T. Chow, Handbook of Applied Hydrology, McGraw-Hill Book Company, New York, 1964.
- 3. T. W. Lambe and R. V. Whitman, Soil Mechanics, John Wiley &
Sons, Inc., New York, 1969.
- 4. M. E. Harr, Groundwater and Seepage, McGraw-Hill Book Company, New York, 1962.
I I 6 Final I
SARGENT& LUNDY ENGINECO@ CHICACO SHAKE SHAKE (Soil Layer Properties and Response / Earthquake) computes response in a horizontally layered semi-infinite system subjected to vertically traveling shear waves. Strain-compatible soil properties are computed within the program. Earthquake motion can be specified at any level of the soil profile and a resulting motion can be computed anywhere else in the profile. The method is based on the continuous solution of the shear wave equation. For soil liquefaction studies, plots of stress time histories at various levels in a soil I profile can also be obtained. The input for the program includes data for the soil profile, curves of strain vs. shear moduli and damping ratios, and the input earthquake motion. The output includes the strain-compatible soil properties, response spectra of object and computed motions and printer and CALCOMP plots of time histories, Fourier spectra and response spectra. Stress time history plots are also included. SHAKE was originally developed by P. B. Schnabel and J. Lysmer of the University of California, Berkeley [1]. It was modified and has been maintained by sargent & Lundy since October 1972 on the UNIVAC 1100 series hardware operating under EXEC 8. 1
SARGENT & LUNDY W ENGlNEERS C HIC AGO To verify the SIIAKE program, the results from SilAKE and the public domain program QUAD 4 [2] were compared for a typical problem. QUAD 4, a finite element program, uses a step-by-step integration technique in the time domain to solve the two-dimensional discrete equations of motion; SIIAKE uses a numerical solution in the frequency domain to solve the one-dimensional wave equation. For the comparison, it was necessary to impose suitable boundary conditions on the finite model for the QUAD 4 analysis to ensure only one-dimensional wave propaga-tion. The problem solved by SilAKE and compared with the QUAD 4 results analyzed the seismic response of a 100 ft layer of dense sand (Figure 1). The properties of the sand are given as: Total unit weight = 125 pcf (K2I mx = 65 Kg = 0.5 The parameter (K2 } max re a es de maximum shear modulus, G max' and effective mean pressure at any depth,y, below the surface. G max " 2) max #m I where (1 + 2K ) o; = 3 of K = Coefficient of lateral pressure at rest of = Effective vertical pressure at depth y. E 2
SARGENT & LUNDY ENGINEERS CHICOGO Damping values and the variation of modulus values with strain were based on published data for sands [3]. I The response of the sand layer was evaluated using the time ) history of accelerations recorded at Taft, California during 1 the 1952 Kern County earthquake as base excitation. The ordi-nates of this time history were adjusted to provide a maximum acceleration of 0.15 g. The maximum shear stresses and accelerations from SHAKE and QUAD 4 are compared in Figure 2 and the response spectra of the surface motions are compared in Figure 3. As illustrated in these figures the two solutions compare favorably. i !I I I 3
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F I I SARGENT & LUNDY EMGINEERS CHICOGO l I REFERENCES
- 1. P. B. Schnabel and J. Lysmer, " SHAKE: A Computer Program for Earthquake Response Analysis of Horizontally Layered Sites," Report No. EERC 72-12, Earthquake Engineering Research Center, Univ. of California, Berkeley, December 1972.
- 2. I. M. Idriss, et al., " QUAD-4, A Computer Program for Evaluating the Seismic Response of Soil Structures by Variable Damping Finite Eleroent," Report No. EERC 73-16, Earthquake Engineering Research Center, University of California, Berkeley, July 1973.
- 3. H. B. Seed and I. M. Idriss, " Soil Moduli and Damping Factors for Dynamic Response Analyses," Report No.
EERC 70-10, Earthquake Engineering Research Center, University of California, Berkeley, December 1970. t I t 7 Final m
e CARGENT O LUNDY ENGINEERO C HIC A t@ SLOPE SLOPE (Slope Stability Analysis) utilizes the theory of equilibrium of forces to determine the factor of safety against sliding of any embankment or slope. It contains the Bishop, Fellenius, and Morgenstern-Price methods of two dimensional stability analysis. In the Bishop and Fellenius mathods, the factor of safety against failure is estimated along a circular surface of failure, whereas any arbitrary failure surface may be chosen for the Morgenstern-Price method. The input includes the slope geometry, soil profile, soil pro-perties (density, cohesion, and the friction angle) and the piezometric surface (s) . The program also has the capability to introduce an earthquake loading assumed as a horizontal gravitational force. Once the problem is input, several execution commands can be used to determine the factor of safety by the various methods. Also, different steges such as end-of-construction, full-lake and sudden-drawdown, can be considered in a single run. The output includes factors of safety for each trial surface and a printer plot of the slope cross section having slope profile, soil profile, water table conditions, and failure surface for the minimum factor of safety. SLOPE was developed and put under ICES (Integrated Civil En-gineering Systems) by William A. Bailey at the Massachusetts Institute of Technology. It has been in the public domain since 1967. Sargent & Lundy currently uses the SLOPE version l 1
SARGENT & LUNDY ENGINEER 3 CHICAGO I maintained by the McDonnell Douglas Automation Company on
.i IBM 370 Series hardware.
I I 1 l 2 LI 9
SARGENT & LUNDY I ENG1NEERS CHICMO REFERENCE 1
" ICES SLOPE - Slope Stability Analysis System", 14cDonnell Douglas Automation Company, 1974.
i l l l t I I l 3 l Final l l
SPRAT SPRAT (Spillway Rating and Flood Routing) was developed by the U.S. Army Corps of Engineers and is in the public domain (HEC program 22-J2-6210, October 1966). The main purpose of the program is to compute a spillway rating curve for a concrete ogee spillway with vertical walls for an. assumed design head, and then make a flood routing of the spillway design flood to determine the maximum water surface. Exact procedures used and other materials referenced as a basis for the program are given in detail in the SPRAT User's Manual supplied by the Corps and reproduced by Sargent & Lundy. Version 3.0 of SPRAT represents a modification of the original version sent by the Corps. Except for one correction, the modifications were made as add-on features to improve the usefulness of the program at Sargent & Lundy, and to improve output readability. SPRAT is currently maintained by Sargent & Lundy on UNIVAC 1100 series hardware, operating under EXEC 8. Since SPRAT is in the public domain, it has been verified by executing a sample problem supplied by the Corps of Engineers, and by comoaring the output produced at Sargent & Lundy to that supplied by the Corps. The input and output from both the Corps' original program and Sargent & Lundy's 3.0 version are numerically identical. The SPRAT program was received from the Corps in August 1970, along with a user's manua.1 dated October 1966, which contained sample input and output listings for a I number of problems. By 1970, the Corps had updated the program, which then produced slightly different results than those published in 1966. This validation compares the results of the sample problem run at Sargent & Lundy to that furnished by the Corps to Sargent & Lundy in 1970. The program contain; one major feature added by Sargent & Lundy: the option to include a second or auxiliary spillway in the rating calculations. This spillway functions only as a broad-crested weir with variable side slopes, and its rating is computed in the same manner as for a service spillway of the same type. l l I I -
I WASP 77 WASP 77 (the Water Surf ace Profiles program) was developed by the U.S. Amy Corps of Engineers at the Hydrologic Engineering Center, Davis, California. (TheCorps program number is 723-X6-L202A, and is commonly referred to as HEC-2). The purpose of the program is to compute and plot (on the printer) water surf ace profiles for river channels of variable cross section, for either subcritical of supercritical flow conditions. The effects of various hydraulic structures, such as bridges, culverts, weirs, embankments, and dams, may be considered in the I computation. The program is used principally for determining profiles for various frequency floods for both natural and modified conditions. The latter may include channel improvements, levees, and floodways. WASP 77 represents by the Corps, the latest and includes the version of the Water latest updates Surface sent by Profile the Corps (M$rNram produced d ication 53 and Error Correction 02), dated October 1; 1977.' WASP 77 is currently maintained on Sargent & Lundy's UNIVAC 1100 series hardware operating under EXEC 8. Since the program is in the public domain, WASP 77 has been validated by running 'g sample problems supplied with the program by the Corps of Engineers, and by -E comparing the results obtained to those published by the Corps. The sample problems supplied by the Corps are documented in their HEC-2 Programmer's Manual 723-02A, dated November 1976, and their program Modification 53, Error Correction 02 letter of October 1,1977, which also contains revised output for sample
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problem 14 (used here for program validation). Two probl m s supplied by the Corps have been run: TEST 5 and TEST 14. In TEST 5, backwater computations are performed on a river up to the point at which it splits into two tributaries. At that point, and starting with the computed water surf ace elevation there, computation continues first up one tributary to its end, and then up the second tributary to its last section. Three profiles are computed in this problem, since three different starting I elevations and flows were specified on input. Summary printouts at each channel section are identical for the Sargent & Lundy run and for the Corps of Engineers run. In TEST 14, .a floodway analysis is performed on a creek by running a 100-year flood through the natural stream channel first, and then running six different encroachment methods to compare the effects against the natural state. A summary of the results of this seven-profile run on a section by section basis is found to be the same for the Sargent & Lundy run and for the Corps of Engineers run. t /}}