ML14015A105

Clinton, Unit 1, Updated Safety Analysis Report, Revision 16, Chapter 2 - Site Characteristics

ML14015A105
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CPS/USAR CHAPTER 02 2.3-1 REV. 11, JANUARY 2005

2.3 METEOROLOGY

2.3.1 Regional Climatology This section provides a description of the general climate of the Clinton Power Station (CPS) site region, as well as the regional meteorological conditions used for design and operating-basis considerations. A climatological summary of normal and extreme values of several meteorological parameters is presented for the first-order National Weather Service Stations at Peoria, Illinois and Springfield, Illinois. Further information regarding the regional climatology was derived from pertinent documents which are referenced in the text. 2.3.1.1 General Climate The CPS site is located near the geographical center of Illinois, approximately 55 miles southeast of the first-order National Weather Service Station at Peoria, Illinois, and 49 miles east-northeast of the first-order National Weather Service Station at Springfield, Illinois.

General climatological data for the region were obtained from U.S. Environmental Science Services Administration (ESSA) Climate of Illinois report (Reference 1), and from the Local Climatological Data Annual summaries for the first-order weather stations at Peoria (Reference 2), and Springfield (Reference 3). The climatic data from Peoria and Springfield are considered to be representative of the climate at the CPS site. The climate of central Illinois is typically continental, with cold winters, warm summers, and frequent short-period fluctuations in temperature, humidity, cloudiness, and wind direction. The great variability in central Illinois climate is due to its location in a confluence zone (particularly during the cooler months) between different air masses (Reference 4). The specific air masses which affect central Illinois include maritime tropical air which originates in the Gulf of Mexico; continental tropical air which originates in Mexico and the southern Rockies; Pacific air which originates in Mexico and eastern North Pacific Ocean; and continental polar and continental arctic air which originates in Canada. As these air masses migrate from their source regions they may undergo substantial modification in their characteristics. Monthly streamline analyses of resultant surface winds suggest that air reaching central Illinois most frequently originates over the Gulf of Mexico from April through August, over the southeastern United States from September through November, and over both the Pacific Ocean and the Gulf of Mexico from December through March (Reference 4). The major factors controlling the frequency and variation of weather types in central Illinois are distinctly different during two separate periods of the year. During the fall, winter, and spring months, the frequency and variation of weather types is determined by the movement of synoptic-scale storm systems which commonly follow paths along a major confluence zone between air masses, which is usually oriented from southwest to northeast through the region. The confluence zone normally shifts in latitude during this period, ranging in position from the central states to the United States-Canadian border. The average frequency of passage of storm systems along this zone is about once every 5 to 8 days. The storm systems are most frequent during the winter and spring months, causing a maximum of cloudiness during these seasons. Winter is characterized by alternating periods of steady precipitation (rain, freezing rain, sleet, or snow) and periods of clear, crisp, and cold weather.

Springtime precipitation is primarily showery in nature. The frequent passage of storm systems, presence of high winds aloft, and frequent occurrence of unstable conditions caused by the close proximity of warm, moist air masses to cold and dry air masses result in this season's CPS/USAR CHAPTER 02 2.3-2 REV. 11, JANUARY 2005 relatively high frequency of thunderstorms. These thunderstorms on occasion are the source for hail, damaging winds, and tornadoes. Although synoptic-scale storm systems also occur during the fall months, their frequency of occurrence is less than in winter or spring. Periods of pleasant, dry weather characterize this season which ends rather abruptly with the returning storminess that usually begins in November. In contrast, weather during the summer months is characterized by weaker storm systems which tend to pass to the north of Illinois. A major confluence zone is not present in the region, and the region's weather is characterized by much sunshine interspersed with thunderstorm situations. Showers and thunderstorms are usually of the air mass type, although occasional outbreaks of cold air bring precipitation and weather typical of that associated with the fronts

and storm systems of the spring months. When southeast and easterly winds are present in central Illinois, they usually bring mild and wet weather. Southerly winds are warm and showery, westerly winds are dry with moderate temperatures, and winds from the northwest and north are cool and dry. The prevailing wind is southerly at both Peoria and Springfield. The frequency of winds from other directions is relatively well distributed. The monthly average wind speed is lowest during late summer at both stations, with the prevailing direction from the south at Peoria and the south-southwest at Springfield. The monthly average wind speed is highest during late winter and early spring at both stations, with the prevailing directions from the west-northwest and the south at Peoria, and the northwest and south at Springfield. Table 2.3-1 presents a summary of climatological data from meteorological stations surrounding the CPS site. The annual average temperature is 50.8

° F at Peoria and 52.7

° F at Springfield. Monthly average temperatures in the CPS site region range from the middle twenties in January to the middle seventies in July. Extreme temperatures in the region range from a maximum of 103° F (Peoria) and 112

° F (Springfield) to a minimum of -20

° F (Peoria) and -22

° F (Springfield). Maximum temperatures in the CPS site region equal or exceed 90

° F with an average of from 17 to 28 times per year. Minimum temperatures in this region are less than or equal to 32

° F for an average of from 119 to 132 times per year (References 2 and 3). Humidity varies with wind direction, being lower with west or northwest winds and higher with east or south winds. The early morning relative humidity is highest during the late summer, with an average of 87% at both Peoria and Springfield. The relative humidity is highest throughout the day during December, ranging from 83% in early morning to 72% at noon at both Peoria and Springfield. Heavy fog with visibility less than 1/4 mile is rare, having an average occurrence of 21 times per year at Peoria and 18 times per year at Springfield. It occurs most frequently during the winter months (References 2 and 3). Annual precipitation at the CPS site area averages about 35 inches per year. For the 40-year period (1937-1976) the minimum annual precipitation was 23.99 inches at Peoria (1940), and 22.88 inches at Springfield (1940). For the same period, the maximum annual precipitation was 50.22 inches at Peoria (1973), and 44.72 inches at Springfield (1941). On the average, about 45% of the annual precipitation occurs in the 4 months of April through August in the CPS site region. However, no month in this region averages less than 4% of the annual total. Monthly precipitation totals have ranged from 13.09 inches (Peoria) to 0.03 inches (Peoria). The maximum 24-hour precipitation at either station was 5.52 inches, recorded at Peoria in May 1927. Snowfall commonly occurs from November through March, with an annual average of 23.4 inches at Peoria, and 22.3 inches at Springfield. The monthly maximum snowfall of 18.9 CPS/USAR CHAPTER 02 2.3-3 REV. 11, JANUARY 2005 inches at Peoria, and 22.7 inches at Springfield, occurred in December 1973. The 24-hour maximum snowfall, which also occurred in December 1973, was 10.2 inches at Peoria, and 10.9 inches at Springfield (Reference 5). The terrain in central Illinois is relatively flat and differences in elevation have no significant influence on the general climate. However, the low hills and river valleys that do exist exert a small effect upon nocturnal wind drainage patterns and fog frequency. 2.3.1.2 Regional Meteorological Conditions for Design and Operating Bases 2.3.1.2.1 Thunderstorms, Hail, and Lightning Thunderstorms occur on an average of 49 days per year at Peoria (1944-1976), and 50 days per year at Springfield (1948-1976)

(References 2 and 3). Approximately 41% of the annual precipitation in the Clinton area falls during thunderstorms (Reference 6). Thunderstorms occur most frequently during the months of June and July; 9 and 8 days per month respectively at Peoria, and 8 and 9 days per month respectively at Springfield. Peoria and Springfield average 5 or more thunderstorm days per month throughout the season from April through September. Both stations average one or less thunderstorm days per month from November through February (References 2 and 3). A thunderstorm day is recorded only if thunder is heard. The observation is independent of whether or not rain and/or lightning are observed concurrent with the thunder (Reference 7). A severe thunderstorm is defined by the National Severe Storms Forecast Center (NSSFC) of the National Weather Service as a thunderstorm that possesses one or more of the following characteristics (Reference 8): a. winds of 50 knots or more, b. hail 3/4-inch or more diameter, and

c. cumulonimbus cloud favorable to tornado formation. Although the National Weather Service does not publish records of severe thunderstorms, the above referenced report of the NSSFC gives values for the total number of hail reports 3/4 inch or greater, winds of 50 knots or greater, and the number of tornadoes for the period 1955-1967 by 1° squares (latitude by longitude). The report shows that during this 13-year period the 1

° square containing the CPS site had 15 hailstorms producing hail 3/4-inch in diameter or greater, 26 occurrences of winds of 50 knots or greater, and 42 tornadoes. At least 1 day of hail is observed per year over approximately 90% of Illinois, with the average number of hail days at a point varying from 1 to 4 (Reference 9). Considerable year-to-year variation in the number of hail days is seen to occur; annual extremes at a point vary from no hail in certain years to as many as 14 hail days in other years. About 80% of the hail days occur from March through August with spring (March through May) being the primary period of occurrence. In the CPS site region, an average of about 22 hail days per 10-year period occurs, with about 55% of all hail days occurring in the spring (Reference 9). Total hailstorm life at a point averages about 7 minutes, with maximum storm life reported as generally not over 20 minutes for Illinois (Reference 6).

CPS/USAR CHAPTER 02 2.3-4 REV. 11, JANUARY 2005 The frequency of lightning flashes per thunderstorm day over a specific area can be estimated by using a formula given by J. L. Marshall (Reference 10), taking into account the distance of the location from the equator:

N = (0.1 + 0.35 sin ) (0.40

+/- 0.20) where: N = number of flashes to earth per thunderstorm day per km 2, and = geographical latitude. For the Clinton Power Station, which is located at 40

° 10'19.5" north latitude, the frequency of lightning flashes (N) ranges from 0.065 to 0.195 flashes per thunderstorm day per km

2. The value 0.195 is used as the most conservative estimate of lightning frequency in the calculations

that follow. Taking the annual average number of thunderstorm days in the site region as 50 (at Springfield), the mean frequency of lightning flashes per km 2 per year is 9.8 as calculated below: yeardaysrmthundersto 50kmdayrmthunderstoflashes195.0 2* = yearkmflashes8.9 2* The area of the CPS site is approximately 14,000 acres, or about 56.7 km

2. Hence the expected frequency of lightning flashes at the site per year is 555, as calculated below:

yearflashes555km7.56yearkmflashes8.9 22=* The exclusion area at the CPS site has a radius of 975 meters (3199 feet), or an area of about

3.0 km2. Hence the expected frequency of lightning flashes in the exclusion area per year is 29, as calculated below:

yearflashes29km0.3yearkmflashes8.9 22=* 2.3.1.2.2 Tornadoes and Severe Winds Illinois ranks eighth in the United States in average annual number of tornadoes. An average of ten tornadoes per year occur on 5 days, based on the period-of-record 1916-1969 (Reference 11). The majority of Illinois tornadoes (65%) occur during the months of March through June.

The statewide probability of a tornado occurrence is greatest during the 7-day period of April 15-

21. Tornadoes can occur at any hour of the day but are more common during the afternoon and CPS/USAR CHAPTER 02 2.3-5 REV. 11, JANUARY 2005 evening hours. About 50% of Illinois tornadoes travel from the southwest to northeast. Slightly over 80% exhibit directions of movement toward the northeast through east. Fewer than 2% move from a direction with some easterly component (Reference 11). Figure 2.3-1 presents the total number of tornadoes (1916-1969) for each county in Illinois. There were 36 tornadoes in the period which originated in the five-county areas (DeWitt, McLean, Logan, Macon, and Piatt) surrounding and including the CPS site. Three tornadoes originated in De Witt County during the 54-year period. The likelihood of a given point being struck by a tornado can be calculated by using a method developed by H. C. S. Thom (Reference 12). Thom presents a map of the continental United States showing the mean annual frequency of occurrence of tornadoes for each 1

° square (latitude by longitude) for the period 1953-1962. For the 1

° square (3634 mi 2 in area) containing the CPS site, Thom computed an annual average of 1.9 tornadoes. Assuming 2.82 mi 2 is the average tornado path area (Reference 12), the mean probability of a tornado occurring at any point within the 1

° square containing the CPS site in any given year is calculated to be .0015. This converts to a mean recurrence interval of 680 years. Using the same annual frequency but an average area of tornado coverage of 3.5 mi 2 (from Wilson and Changnon, Reference 11), the mean probability of a tornado occurrence is .0018. More recent data (Reference 8) containing tornado frequencies for the period 1955-1967 indicate an annual tornado frequency of 3.2 for the 1

° square containing the CPS site. This frequency, with Wilson and Changnon's average path area of 3.5 mi 2, results in an estimated mean tornado probability of .0031, with a corresponding mean return period of 325 years. The above results were presented in order to provide a reasonable estimate of tornado probability without addressing the accuracy of the estimate. Because of the uncertainties in regard to tornado frequency and path area data, the annual tornado probability for the CPS site area is best expressed as being in the range of .0015 to .0030, with a mean tornado return period of 330 to 670 years. A conservatively high estimate can be taken to be .0031, with a corresponding mean return period of 325 years. The following are the design-basis tornado parameters (Reference 13) that were used for the Clinton Power Station: a. rotational velocity = 290 mph b. maximum translational velocity = 70 mph

c. radius of maximum rotational velocity = 150 ft

d. pressure drop = 3.0 psi e. rate of pressure drop = 2.0 psi/sec A design wind velocity of 85 mph (100-year recurrence interval) was used in the design of Clinton Power Station Seismic Category I structures. This design wind velocity is estimated from the analysis presented in Figure 2 of the "American National Standard Building Code Requirements for Minimum Design Loads in Buildings" (Reference 14). The vertical velocity distribution and gust factors employed for the design wind loading are those specified in Reference 14 for exposure type C (see Subsection 3.3.1).

CPS/USAR CHAPTER 02 2.3-6 REV. 11, JANUARY 2005 2.3.1.2.3 Heavy Snow and Severe Glaze Storms Severe winter storms, which usually produce snowfall in excess of 6 inches and are often accompanied by damaging glaze, are responsible for more damage in Illinois than any other form of severe weather, including hail, tornadoes, or lightning (Reference 15). These storms occur on an average of five times per year in the state. The state probability for one or more severe winter storms in a year is virtually 100% while the state probability for three or more in a year is 87%. A typical storm has a median point duration of 14.2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. Point durations have ranged from 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> to 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> during the 61-year period-of-record 1900 to 1960 used in the severe winter storm statistical analyses (Reference 15). Data on the average areal extent of severe winter storms in Illinois show that they deposit at least 1 inch of snow over 32,305 mi2, with more than 6 inches of snow covering 7500 mi

2. Central Illinois (including the CPS site) had 107 occurrences of a 6-inch snow or glaze damage area during the years 1900-1960. About 42 of those storms deposited more than 6 inches of snowfall in De Witt County (Reference 15). In the Springfield area, the maximum 24-hour snowfall was 15.0 inches, and the maximum monthly snowfall was 24.4 inches, both of which occurred in February 1900. On the average, heavy snows of 4 to 6 inches have occurred one to two times per year (Reference 15). The 2-day and 7-day maximum snowfall values for selected recurrence intervals in the Clinton Power Station area are as follows (Reference 15):

2-yr 5-yr 10-yr 20-yr 30-yr 50-yr 2-day 7.0 8.6 10.2 12.1 13.4 15.2 7-day 7.6 10.1 12.8 16.3 18.7 22.0 The listed value is the number of inches of snowfall which would be equalled or exceeded in the given interval of years. Sleet or freezing rain occurs during the colder months of the year when rain falls through a shallow layer of cold air with a temperature below 32

° F from an overlying warm layer of temperature above 32

° F. The rain becomes supercooled as it descends through the cold air. If it cools enough to freeze in the air, it descends to the ground as sleet; otherwise, it freezes upon contact with the ground or other objects, causing glaze. In Illinois, severe glaze storms occur on an average of about three times every 2 years. Statewide statistics indicate that during the 61-year period 1900-1960, there were 92 glaze storms defined either by the occurrence of glaze damage or by occurrence of glaze over at least 10% of Illinois. These 92 glaze storms represent 30% of the total winter storms in the period. The greatest number of glaze storms in 1 year was six (1951); in 2 years, nine (1950-1951); in 3 years, ten (1950-1952); and in 5 years, fifteen (1948-1952). In an analysis of these 92 glaze storms, Changnon (Reference 15) determined that in 66 storms, the heaviest glaze disappeared within 2 days; in 11 storms, 3 to 5 days; in 8 storms, 6 to 8 days; in 4 storms, 9 to 11 days; and in 3 storms, 12 to 15 days. Fifteen days was the maximum persistence of glaze. Within the central third of Illinois, eleven localized areas received damaging glaze in an average 10-year period; the CPS site area averages slightly over 5 days of glaze per year (Reference 15). Ice measurements recorded in some of the most severe Illinois glaze storms are shown in Table 2.3-2 (Reference 15). The listing reveals that severe glaze storms depositing ice of moderate to CPS/USAR CHAPTER 02 2.3-7 REV. 11, JANUARY 2005 large radial thickness may occur in any part of Illinois. An average of one storm every 3 years will produce glaze ice 0.75 inch or thicker on wires (Reference 15). Strong winds during and after a glaze storm greatly increase the amount of damage to trees and power lines. In studying wind effects on glaze-loaded wires, the Association of American Railroads (Reference 16) concluded that maximum wind gusts were not as significant (harmful)

a measure of wind damage as were speeds sustained over 5-minute periods. Moderate wind speeds (10-24 mph) occurring after glaze storms are most prevalent. Wind speeds of 25 mph or higher are not unusual however, and there have been 5-minute winds in excess of 40 mph with glaze thicknesses of 0.25 inch or more (Reference 15). Table 2.3-3 presents specific glaze thickness data for the five fastest 5-minute speeds and the speeds with the five greatest measured glazed thickness for 148 glaze storms throughout the country during the period 1926-1937 (Reference 15). Although these data were collected from various locations throughout the United States, they are considered applicable designed values for locations in Illinois. The roofs of safety-related structures are designed to withstand the snow and ice loads due to a winter probable maximum precipitation (PMP) with a 100-year recurrence interval antecedent snowpack. A 100-year return period snowpack of 22 psf (or 22 inches of snowpack) was obtained from the American National Standards building code requirements (Reference 14);

however, for design a 100-year return period snowpack of 25 psf is used. The weight of the accumulation of the winter PMP from a single storm is 79 psf (15.2 inches of precipitable water, or about 152 inches of fresh snow), which was taken as the 48 hour5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> PMP

during the winter months (December through March) (Reference 18). Thus the weight of snow and ice on the roof of each safety-related struct ure can be conservatively estimated as 104 psf. 2.3.1.2.4 Ultimate Heat Sink Design The meteorological conditions used in evaluation of the performance of the ultimate heat sink were obtained from Peoria, Illinois meterological data for the period of record 1949 through 1971, which was supplemented with meterological data from Springfield, Illinois for the period January 1, 1952 through December 31, 1956. The critical period which showed the maximum station intake temperature was July 1, 1964 through September 30, 1964. The mean values of wind speed, dry bulb temperature, wet bulb temperature, and dew point temperature for this 92-day period were 8.4 mph, 71

° F, 64° F, and 59

° F, respectively. This period was also found to be a period of high evaporative losses. For further details, see Subsection 9.2.5. 2.3.1.2.5 Inversions and High Air Pollution Potential Weather records from many U.S. weather stations have been analyzed by Hosler (Reference 19) and Holzworth (Reference 20) with the objective of characterizing atmospheric dispersion potential. The seasonal frequencies of inversions based below 500 feet for the CPS site are shown by Hosler (Reference 19) as:

CPS/USAR CHAPTER 02 2.3-8 REV. 11, JANUARY 2005 INVERSIONS BELOW 500 FEET SEASON % OF TOTAL HOURS % OF 24-HOUR PERIODS WITH AT LEAST 1 HOUR OF INVERSION Winter 29 53 Spring 29 67 Summer 33 81 Fall 39 8 Since central Illinois has a primarily continental climate, inversion frequencies are closely related to the diurnal cycle. The less frequent occurrence of storms in summer and early fall produces a larger frequency of nights with short-duration inversion conditions. Holzworth's data give estimates of the average depth of vigorous vertical mixing, which give an indication of the vertical depth of atmosphere available for mixing and dispersion of effluents. For the CPS region, the seasonal values of the mean daily mixing depths are (Reference 20):

MEAN DAILY MIXING DEPTHS (meters) SEASON MORNING AFTERNOON Winter 400 690 Spring 490 1500 Summer 330 1600 Fall 390 1200 When daytime (maximum) mixing depths are shallow, pollution potential is highest. Holzworth has also presented statistics on the frequency of episodes of high air pollution potential, defined as a combination of low mixing depth and light winds (Reference 20). Holzworth's data indicate that during the 5-year period 1960-1964, the region including the CPS site experienced no episodes of 2 days or longer with mixing depths less than 500 meters and winds less than 2 meters per second. There were two such episodes with winds remaining less than 4 meters per second. For mixing heights less than 1000 meters and winds less than 4 meters per second, there were about nine episodes in the 5-year period lasting 2 days or more but no episodes lasting 5 days or more. Holzworth's data indicate that central Illinois is in a relatively favorable dispersion regime with respect to low frequency of extended periods of high air pollution potential. 2.3.2 Local Meteorology The onsite meteorological monitoring program began at the Clinton Power Station (CPS) site on April 13, 1972. Onsite meteorological instrumentation is described in Subsection 2.3.3. Data from this installation have been used in preparation of the local meteorological summaries. These data are from a 5-year period of record (April 13, 1972 through April 30, 1977) and, therefore, can be considered representative of long-term site meteorology.

CPS/USAR CHAPTER 02 2.3-9 REV. 11, JANUARY 2005 2.3.2.1 Normal and Extreme Values of Meteorological Parameters 2.3.2.1.1 Wind Summaries Detailed wind records, suitable for the preparation of wind roses, are available from the plant site for April 1972 through April 1977. Monthly and period of record wind roses were constructed for the 33-foot (10 meter) level of the onsite meteorological tower. The period of record wind rose is presented in Figure 2.3-2. The composite monthly wind roses are found in Figures 2.3-3 through 2.3-14. Seasonal variations are evident from monthly data. Winds from the sector SSE through WNW tend to dominate in most months. Winter months show generally higher wind speeds, fewer calms and more WNW winds than do the summer months. For the period of record, the following frequencies of occurrence were observed at the CPS site for the specified wind speed intervals: Wind Speed Percent of Occurence

< 0.3 mps (calm) 0.3% 0.3 to 1.4 mps 7.7% 1.5 to 3.0 mps 28.2% 3.1 to 5.0 mps 30.7% 5.1 to 8.0 mps 23.7% > 8.0 mps 9.4% There were two occurrences of persistence of wind direction for 33 hours3.819444e-4 days <br />0.00917 hours <br />5.456349e-5 weeks <br />1.25565e-5 months <br /> (the longest persistence observed). These occurred in two sectors, the SSW and the NE. 2.3.2.1.2 Temperatures Temperatures at the Clinton Power Station meteorological monitoring site were measured at the 10 meter level of the tower. The average daily temperature for the period of record is 10.5

° C (50.9° F). The absolute maximum is 35.2

° C (95.4° F) and the absolute minimum is -28.8

° C (-19.8° F). Period of record and composite monthly summaries of onsite temperature data are presented in Tables 2.3-4 through 2.3-6. 2.3.2.1.3 Atmospheric Moisture 2.3.2.1.3.1 Relative Humidity The relative humidity for a given moisture content of the air is inversely proportional to the temperature cycle. A maximum relative humidity usually occurs during the early morning hours, and a minimum is typically observed in the midafternoon. For the annual cycle, the lowest humidities occur in midspring, the winter months experience the highest. Table 2.3-7 presents a relative humidity summary for the Clinton Power Station.

CPS/USAR CHAPTER 02 2.3-10 REV. 11, JANUARY 2005 2.3.2.1.3.2 Wet Bulb The wet bulb temperature is not as strong a function of the ambient temperature as the relative humidity and is used for evaporative cooling system modeling studies. The wet bulb temperature is defined to be the temperature to which an air parcel may be cooled by evaporating water into it at constant pressure until it is saturated. All latent heat utilized in the process is supplied by the air parcel. Summaries of wet bulb temperatures are presented in Table 2.3-8. These values are calculated from the dew point and ambient temperatures, assuming a constant standard sea level pressure of 1013.25 millibars. 2.3.2.1.3.3 Dew Point Temperature Dew point temperature is a measure of absolute humidity in the air. It is the temperature at which the air must be cooled to cause condensation to occur, assuming pressure and water

vapor content remain constant. Composite monthly and period of record dewpoint summaries are presented in Tables 2.3-9 through 2.3-11. 2.3.2.1.3.4 Precipitation The average yearly precipitation for the period of record for the Clinton Power Station site is 25.47 inches. Period of record and composite monthly precipitation data appear in Table 2.3-12. The months of March and June are the wettest, and December, January, and February are

the driest. 2.3.2.1.3.5 Fog Fog is an aggregate of minute water droplets suspended in the atmosphere near the surface of the earth. According to international definition, fog reduces visibility to less than 0.62 mile. According to United States observing practice, ground fog is a fog that hides less than 0.6 of the sky, and does not extend to the base of any clouds that may lie above it. Ice fog is fog composed of suspended particles of ice. It usually occurs in high latitudes in calm clear weather at temperatures below -20

° F and increases in frequency as temperature decreases (Reference 21). Since local data are not available to assess the fog statistics at Clinton, data are presented for Springfield, Illinois and Peoria, Illinois. Fog is a very local phenomenon; thus, these data should be considered only as regional estimates. The average number of days during which heavy fog (visibility less than 1/4 mile) occurs is as follows (Reference 22). SPRINGFIELD Peoria January 2 3 February 3 3 March 2 2 April 1 1 May 1 1 June 1/2 1

CPS/USAR CHAPTER 02 2.3-11 REV. 11, JANUARY 2005 SPRINGFIELD Peoria July 1 1 August 1 1 September 1 1 October 1 1 November 2 2 December 3 3 Year 18.5 20 The yearly average is 18.5 days at Springfield and 20 days at Peoria; the highest occurrence of fog at both locations is in the winter months. Tables 2.3-13 and 2.3-14 summarize the occurrence of all fog for Peoria and Springfield, respectively. These summaries were prepared by processing the digital data tapes for these stations. Fog extracted from these tapes were any of the three fogs coded "fog," "ground fog,"

and "ice fog" which occurred in column 132, "obstruction to vision," on the Airways Surface

Observations tapes. The percentage of the total observations that fog was reported for Peoria and Springfield is given in the first column of Tables 2.3-13 and 2.3-14. The hour and the percentage of observations for that hour of the maximum and minimum fog occurrence are given in the next four columns. Peoria shows a higher frequency of fog in all months than Springfield. The long-term annual average percent of hourly observations with any intensity of fog for Peoria and Springfield are 11.3% and 9.1%, respectively. The occurrence of prolonged periods of fog is also greater for Peoria. Although information on fog is generally a very local phenomenon, the expected occurrences at the Clinton Power Station should be within the range represented by these two

stations. 2.3.2.1.4 Atmospheric Stability For estimates of average dispersion over extended periods, the joint probability of occurrence of wind speed, wind direction, and atmospheric stability must be known. These probabilities, or frequencies, have been generated from onsite data using the vertical temperature gradient and the variability of the horizontal wind to estimate atmospheric stability in accordance with Regulatory Guide 1.23. Summaries of wind speed-wind direction-atmospheric stability joint frequencies appear in Tables 2.3-15 through 2.3-22. The following data summarize the percent frequencies of occurrence for each stability class (determined from the temperature gradient) recorded at the Clinton Power Station site.

CPS/USAR CHAPTER 02 2.3-12 REV. 11, JANUARY 2005 A B C D E F G 4.34 3.58 5.38 40.10 26.52 10.93 8.88 Unstable (A, B, C) 13.30%

Neutral (D) 40.10%

Stable (E, F, G) 46.33%

The combination of E stability and calm winds (< 0.3 mps) occurred 0.06% of the time; F and calm, 0.06%; G and calm, 0.12%. 2.3.2.2 Potential Influence of the Plant and Its Facilities on Local Meteorology Operation of the station will influence the local micrometeorology as a result of discharging warm water into the cooling lake. The principal meteorological effect of this will be to produce a steam fog over the lake when cold air (~41° F or less) moves over the significantly warmer (~59° F or higher) lake water. The rate of condensation of evaporated water vapor (and thus the formation of steam fog) will be greatest at the lower ambient air temperature associated with the winter months. With heavy steam fog and relatively light wind speeds (~2 meters per second - 5 mph - or less), noticeable drift of the steam fog off the lake surface is possible. Icing caused by condensed water vapor from the lake will have a primary effect on vertical surfaces adjacent to the lake shore. Horizontal surfaces will accumulate much less rime. Observations of icing conditions from the Dresden Nuclear Power Station in Illinois indicate that icing on horizontal surfaces is not a significant problem beyond the first 200 feet from the edge

of the lake. 2.3.2.2.1 Topographical Description Figure 2.3-15 is a topographic map of the area within 50 miles of the Clinton Power Station site. Figure 2.3-16 is a topographic map of the areas within 5 miles of the site. Figure 2.3-17 shows topographic cross sections in each of the 16 compass directions radiating from the site. The crosshatched sections represent the areas to be filled in by the creation of the cooling lake. The station is located at an elevation of approximately 735 feet MSL. Within the 5-mile radius, no land elevation is above 760 feet or below 640 feet. Much of this modest relief is due to the shallow valleys surrounding the North Fork of Salt Creek and Salt Creek. These valleys form the boundaries of the Clinton Power Station cooling lake (Lake Clinton). The surface of Lake Clinton is 690 feet. Thus, a large portion of the topographical relief in the immediate area is filled by the lake. Lake Clinton presents a discontinuity in the ground surface over which diffusing gases must travel. The lake presents a temporary smoother surface than the land over which the air parcels travel. Theoretically, this reduces the natural turbulence and thus the resulting diffusion. At the same time, however, the reduced frictional effects will allow a slight increase in the wind speed, thus adding to the rate of diffusion. In view of the relatively short travel distances across the lake for releases from the station under any wind direction, no adjustments in the diffusion calculations are proposed at this time to account for the reduction in surface roughness caused

by the lake.

CPS/USAR CHAPTER 02 2.3-13 REV. 11, JANUARY 2005 A more significant impact of the lake will be the warm surface it presents to the atmosphere which, during nighttime and the winter, will be significantly warmer than the surrounding ground. This increase in temperature will cause the layer of air in contact with the lake to achieve a neutral lapse rate, especially when stable conditions prevail over the land. Thus, material released from a ground-level source would receive additional diffusion in the vertical over the lake than would be computed using a stable delta T stability category determined from the meteorological tower. However, due to the dimensions of the lake and its orientation with respect to the station, no adjustments are proposed at this time to the diffusion calculations.

Any additional dispersion contributed by the lake temperature effects will add to the conservatism of the accident and routine diffusion estimates. The natural topography of the area around the site will not significantly affect the diffusion estimates. 2.3.2.2.2 Prediction of Cooling Lake Steam Fog The cooling lake with once-through cooling provides a source of open water during the winter months. It is possible that cold air passing over the relatively warmer water surface can become saturated with respect to water vapor. When sufficient evaporated water vapor condenses into droplets, steam fog occurs and the transparency of the air is reduced. The characteristics of such steam fog will vary with the water temperature, the distance traveled over the water, and the low-level ambient air temperature, relative humidity, vertical and horizontal stability, and the transporting wind speed.

An analytical model was used that accounts for the processes of evaporation, condensation, and diffusion downwind and includes the variables listed previously as input conditions. A description of the model is provided in Attachment A2.3 (Analytical Fog Model). A portion of the cooling lake will be subjected to increases in water temperature due to the operation of the station. These increases in water temperature were determined by use of the LAKET computer model. The physical characteristics of the lake, such as time-varying temperature and natural and forced evaporation, were predicted by LAKET (Transient Lake Temperature Prediction) (References 23, 24 and 25). This program simulated the effects of varying weather conditions and station heated-water discharge on the surface temperature and evaporation rates of a lake or river. The time-varying temperature distribution along the water body's central axis is computed against time, along with the natural and forced evaporation. In the case of lakes, the variation in the lake level is also computed. Inputs to the computer program include data on the lake, the station, and the weather. Lake data include total surface area, salt content, seepage rate, initial temperature, and the length and width of the segments used in the analysis. Station data include temperature rises, flow rates, latitude, longitude, and altitude. Weather data include dates, wind speed, dry bulb temperatures, relative humidity, dew point, barometric pressure, air vapor pressure, cloud cover, and precipitation. Output from the program provides time-var ying temperature along the water body, natural and forced evaporation, and plots of temperature vs. time at nine locations.

The computational approach consists of modeling the body of water into an idealized system of prismatic volumes, each having geometric and physical characteristics (i.e., width, depth, area, and flow) unique to its location and time. Using inputted weather data, the natural water CPS/USAR CHAPTER 02 2.3-14 REV. 11, JANUARY 2005 temperature is determined, and based on the station rise, the downstream temperatures are computed. The one-dimensional finite-difference procedure discussed in Reference 24 is used. Hydraulic and thermal balances are utilized along with the energy budget method for determining the evaporation from the lake or river. The latter takes into account solar radiation, reflected solar radiation, energy transferred from the lake back to the atmosphere, and other factors. Specific areas of interest within the immediate vicinity of the lake were defined for detailed study and evaluation of the steam fog potential and resulting impact. The seven areas selected are as follows: a. Area 1 - road crossing the lake south of DeWitt, b. Area 2 - the county road that runs east-west along the southern edge of the lake just west of Route 14, c. Area 3 - Route 10 where it runs along the southern edge of the lake,

d. Area 4 - the NW-SE portion of Route 10 that is parallel to the spillway,

e. Area 5 - Route 10 and the connecting roads that run N-S along the western edge of the site, f. Area 6 - that portion of U.S. Route 54 that is close to the lake including the bridge area over the lake, and g. Area 7 - the reactor building complex. Calculations showed no significant probability of the lake steam fog extending to DeWitt. Similarly, the probability of the lake steam fog reaching the town of Lane is so low that the town did not require designation as a special area. The remaining sections of roads around the lake also were not affected significantly by the predicted lake steam fog. The steam fog prediction model described in Attachment A2.3 was used to calculate the occurrence of restricted visibility caused by steam fog in each of the specified areas. This determination required the calculation of evaporation and diffusion for each of six to ten combinations of temperature and relative humidity for each of the seven major wind directions that would affect one or more of the areas of interest. This process was repeated for each month to account for the monthly difference in water temperature. The results were several hundred maps showing the concentration of water vapor and water droplets for the lake and adjacent areas. The time required to run the model and to evaluate the results did not permit complete variation of all the variables that would influence the horizontal extent and intensity of steam fog from the lake. Therefore, a set of values was selected and used in the model to produce what are considered the probable "worst case" for contiguous 30-day periods. Assumptions and variables used in the model are described in Attachment A2.3. Briefly, these assumptions are as follows:

CPS/USAR CHAPTER 02 2.3-15 REV. 11, JANUARY 2005 a. The wind speed shear in the layer into which water is evaporated is 1 m/sec. b. The mean wind speed in the layer is 1 m/sec. c. The vertical and horizontal stability in the layer is Pasquill stability category C during the period August through April. d. The calculated lake water temperatures apply uniformly across the width of the lake. e. The edges of the lake do not freeze and reduce the amount of surface water available for evaporation. f. Visibility is defined by the empirical cu rves derived from previous fog research and presented in Attachment A2.3. g. The horizontal visibility is measured at a height of 1 meter above the surface of the lake. The vast majority of predicted hours of steam fog off the lake occurred when the air temperature was 5° C or less with the water temperature 10

° C to 25° C or higher. These conditions would produce an "unstable" lapse rate within the layer of interest. Stability category C was selected to represent this type of stability lapse rate. The calculated number of hours of various categories of visibility due to steam fog from the lake for each selected area are presented in Tables 2.3-23 through 2.3-29. The values in these tables are the sum of all the hours of various combinations of air temperature, relative humidity, and wind direction that could affect a given area. Thus, the values do not apply uniformly over an entire area, but rather just for that portion that is immediately downwind of the lake for the occurring wind direction. The fog prediction model has been used to predict the hours of steam fog during the summer months. Using the same assumptions as for the other months of the year, the model predicts a greater number of hours than would be expected or is verified by the calibration data. Rather than attempt to refine the model to obtain more precise (smaller) values, the derived values are presented to serve as an upper limit on the number of hours of off-lake steam fog that could be expected. The magnitude of lake steam fog during the summer months was not fully determined. Results showed steam fog forming during the cooler nighttime temperatures and periods of high relative humidity. Additional calibration of the model is required for these summer warm fog conditions before values can be presented. The values presented in Tables 2.3-23 through 2.3-29 are considered to be representative of the worst probable monthly average conditions expected for the month. The basis for this conclusion is the conservative nature of the input values, described in the preceding paragraph,

that were used in the model. Normal station operating conditions (normal station operating conditions are defined as a lake elevation of 690 feet; a 70% load factor for February, March, April, May, October, and November; and an 80% load factor for June, July, August, September, and January) were used CPS/USAR CHAPTER 02 2.3-16 REV. 11, JANUARY 2005 to determine the lake water temperatures. Occasional periods of heavier loads or lower lake elevations would not significantly affect the predicted steam fog hours. This conclusion is possible because continuous (365 days per year) operation under extreme station operating conditions results in a net increase in water temperature of approximately 5.5

° F at the discharge and 3

° F at the intake during the winter. These changes would have their greatest impact immediately below the discharge point; the impact would decrease rapidly downstream with the requirement to transport more or heavier fog farther off the lake surface into the areas of interest. Test results with slightly (2

° F to 5° F) increased water temperatures showed negligible change in the resulting steam fog for a given meteorological condition. Probably the most influential conservative factor used in the model is the assumed low-level wind speed of 1 m/sec. The impact of this assumption is indicated as follows. First, it reduces the thickness of the layer to reach saturation more rapidly and achieve a greater concentration of condensed water vapor. The low wind speed then moves the steam fog off the lake in a relatively uniform mass, albeit a shorter distance, before off-lake evaporation improves the visibility. Wind speed data collected at Clinton 10 meters above the ground during the period of record showed that only 8% of the hours had a wind speed of 1.5 m/sec or less. Nevertheless, it is felt that wind speed of 1 m/sec should be used for a conservative approximation of the near-surface wind speed that will affect the horizontal visibility in the first 3 meters above the ground in the designated areas of interest. The maps (not included) produced by the computer fog model show the horizontal extent of various concentrations of water vapor or condensed water that occur with a given wind direction for a specified combination of air temperature and relative humidity. Analyses of these maps show that the maximum extent of reduced visibility beyond the lake from the lake steam fog will be generally confined to the area that is south of the lake and east of the town of Lane. Steam fog can occasionally drift over U.S. Route 54 where it passes near the northern edge of the lake. A shallow open flume about 300 feet wide will be used to carry the discharge water from the station to the discharge point in the lake approximately 3 miles due east of the station. Because of the water temperature, steaming in the flume is expected with the same frequency as for Area 1 (Table 2.3-23). However, the relative narrow width of the flume will limit the volume of air exposed to the water surface and thereby limit the amount of air to reach saturation. Under most meteorological conditions, any excess water vapor acquired over the flume will mix with the drier ambient air as soon as the parcel of air is beyond the flume. With low-level wind speeds of less than 2 m/sec, the expected extent of significantly reduced visibility due to steam fog from the flume will be limited to, at mo st, a few hundred feet immediately downwind of the flume. With higher wind speeds, any steam fog should dissipate within 200 feet of the flume. A greater horizontal extent will occur when the ambient air is very near saturation prior to exposure to the flume. In this case, natural fog would be expected and the steam fog from the flume would act to increase the intensity of the ambient restriction to visibility immediately

downwind of the flume. The impact of fogging and icing conditions on emergency procedures for a coincident station accident is primarily in the area of transportation. The safe speed of vehicles through the area downwind of the lake and affected by lake steam fog could be reduced if the lake steam fog is sufficiently dense. As a conservative estimate, a speed of 10 to 15 mph could still be maintained through an affected area in all but the most extreme cases.

CPS/USAR CHAPTER 02 2.3-17 REV. 13, JANUARY 2009 The maximum horizontal extent of steam fog from the lake along a road is on the order of 1 mile or less. The extent of extremely dense steam fog would be limited to the road area immediately adjacent to the lake. Once vehicles are through an affected area, the speed of the vehicle is controlled by other factors. Icing from lake steam fog should not be a problem. Roads located 500 feet or more from the lake are not expected to be affected by ice from the lake. Vertical surfaces within 500 feet downwind of the lake could accumulate rime ice under certain meteorological conditions. A horizontal surface, such as a road bed, is seldom affected by lake ice if it is 50 feet or more from the edge of the lake. If significant icing should occur on any critical road due to natural or cooling lake influences, standard highway maintenance procedures will be followed to reduce the impact of the ice on vehicle movement over the affected critical roads. The white or hoary accumulation of ice on vertical surfaces along a roadway can alert drivers and maintenance personnel to the possibility of icing conditions on the road. 2.3.2.3 Local Meteorological Conditions for Design and Operating Bases Design and operating bases such as tornado parameters, glaze thickness, and winter probable maximum precipitation are statistics which by definition and necessity are based upon long-term regional records. While data collected at the Clinton onsite meteorological monitoring system can be considered representative of long-term site meteorology, long-term regional data are most appropriate for use as conservative estimates of climatological extremes. Therefore, all design and operating basis conditions were based upon regional meteorological data, as described in Subsection 2.3.1.2. 2.3.3 Onsite Meteorological Measurements Program The meteorological monitoring program began at the Clinton Power Station site on April 13, 1972. The instrument systems and their locations were selected with emphasis on compliance

with Regulatory Guide 1.23. A tower with two levels of instrumentation was erected. There are no trees, tall obstructions or significant topographical features in the immediate vicinity of the tower. The ground under the tower is covered with short natural grasses and weeds. The location of the tower is shown in Figure 2.3-18. The meteorological measurements program at the Clinton site consists of monitoring wind direction, wind speed, temperature, dewpoint, and precipitation. The Main Tower is instrumented at the 10 meter and 60 meter levels. All parameters are recorded digitally and displayed in the Main Control Room. Data recovery is expected to exceed 90% for all parameters. Two methods of determining atmospheric stability are used: delta T (vertical temperature difference) is the principal method; sigma theta (standard deviation of the horizontal wind direction) is available for use when delta T is not available. These data, referenced in ANSI/ANS 2.5 (1984), are used to determine the meteorological conditions prevailing at the plant site. The meteorological tower is equipped with instrumentation that conforms with the system accuracy recommendations of Regulatory Guide 1.23. The equipment is placed on booms oriented into the generally prevailing wind at the site. Equipment signals are brought to an instrument shack with controlled environmental conditions. The shack at the base of the tower CPS/USAR CHAPTER 02 2.3-18 REV. 13, JANUARY 2009 houses the recording equipment, signal conditioners, etc., used to process and retransmit the data to the end-point users. Recorded meteorological data are used to generate wind roses and provide estimates of airborne concentrations of gaseous effluents and projected offsite radiation dose. In addition to the meteorological instruments, an unused antenna for the Alert and Notification System is mounted on this tower at approximately 170 feet high. Meteorological monitoring instruments have been placed on the microwave tower to act as a backup to the existing meteorological monitoring instruments on the meteorological tower. The microwave tower is 250 feet high with instrumentation (wind speed and direction) installed at the 33-foot (10-meter) level. The current antenna for the Alert and Notification System is mounted on this tower. The location of the tower is shown on Figure 2.3-18. The monitoring panel, located in a shelter at the base of the microwave tower, is a microprocessor based system which is used to collect, process, format and record all the meteorological data supplied. The data is displayed locally and is accessible for review and trending at the 800 foot elevation of the Control Building. Heating and air conditioning are thermostatically controlled in the shelter to provide a controlled environment for the data processing equipment. The NRC requested an additional tape containing weather data from Clinton Power Station (1972-1979). It provided to the NRC under separate cover on September 18, 1981 (Q&R

451.01). In response to a request for a complete record for 12 consecutive months of hour by hour onsite meteorlogical data, the following is provided (Q&R 451.02) (a) The selected period is one year of data from 73/15/00 to 74/14/23. That is January 15, 12:00 A.M., 1973 to January 14, 11:59 P.M., 1974. (Date is YY/Julian Day/HH.) (b) Attachment A gives the dates and hours of missing data in the selected period.

Attachment B provides recommended substitute values for the missing data. (c) The bases for the substitutions were extrapolations and interpolation using data before and after the missing period. There were no lengthy periods of missing data which required more involved methods. There are no recommended values for precipitation given.

CPS/USAR CHAPTER 02 2.3-19 REV. 11, JANUARY 2005 ATTACHMENT A (Q&R 451.02) List of missing data within the period 73/15/99 to 74/14/23 (Dates are YY/Julian Day/HH) DIR = Wind Direction, Degrees RNG = Wind Direction Variability, Degrees SPD = Wind Speed, Meters per second T = 10 Meter Temperature, Degrees Celsius DT = 60m-10m Temperature, Degrees Celsius DP = Dew Point, Degrees Celcius 10 ro 60 meter levels P = Precipitation, inches Date Hours Missing Parameter(s) 73/16/10 1 10m DIR, RNG, SPD 73/32/4 7 60m DIR, RNG, SPD 73/46/00 43 10m DIR, RNG, SPD 73/48/03 61 60m DIR, RNG, SPD 73/51/7 28 60m DIR, RNG, SPD 73/61/6 5 T, T, 10 and 60m DP 73/66/1 4 10m DIR, RNG, SPD 73/94/12 1 T, T, 10 and 60m DP 73/112/16 4 10 and 60m DIR, RNG, SPD, DP, T, T 73/120/21 1 T, T, 10 and 60m DP 73/128/8 8 P 73/135/11 2 T, T, 10 and 60m DP 73/144/19 236 P 73/168/13 46 P 73/179/8 37 60m DIR, RNG, SPD 73/195/9 3 T, T, 10 and 60m DP 73/197/15 2 T, T, 10 and 60m DP 73/211/18 28 10m DIR, RNG, SPD 73/219/10 23 60m DIR, RNG 73/228/18 14 60m DIR, RNG, SPD 73/233/10 26 60m DIR, RNG, SPD 73/244/7 76 60m DIR, RNG 73/265/21 10 10m SPD 73/266/19 5 10m SPD 73/269/9 9 T, T, 10 and 60m DP 73/318/3 32 60m DIR, RNG, SPD 73/319/22 19 60m DIR, RNG, SPD 73/320/17 4 10m SPD CPS/USAR CHAPTER 02 2.3-20 REV. 11, JANUARY 2005 ATTACHMENT A (CONT'D)

(Q&R 451.02)

Date Hours Missing Parameter(s) 73/320/17 91 P 73/333/11 4 60m DIR, RNG, SPD 73/335/19 41 60m DIR, RNG, SPD 73/346/1 10 60m SPD 73/358/21 18 10m DIR, ENG 73/361/3 7 10m SPD 73/10/19 15 10m SPD CPS/USAR CHAPTER 02 2.3-21 REV. 11, JANUARY 2005 ATTACHMENT B (Q&R 451.02) Substitute values for missing data within period from 73/15/00 to 74/14/23 (Dates are YY/Julian Day/HH) DIR = Wind Direction, Degrees RNG = Wind Direction, Variability, Degrees SPD = Wind Speed, Meters per second T = 10 Meter Temperature, Degrees Celsius T = 60m-10m Temperature, Degrees Celsius DP = Dew Point, Degrees Celcius 10 or 60 meter levels P = Precipitation, inches Date Hour Parameter Hour Parameter 73/16 10M DIR. RNG. SPD. 179. 50. 80. 73/32 60M DIR. RNG. SPD. 4 137. 55. 10.0 5 131. 52. 10.1 6 118. 53. 10.3 7 128. 60. 10.5 8 135. 56. 10.4 9 138. 50. 11.0 10 145. 52. 10.9 73/46 10M DIR. RNG. SPD. DIR. RNG. SPD 0 327. 57. 4.4 11 320. 63. 7.4 1 320. 60. 4.0 12 317. 43. 7.4 2 323. 50. 5.0 13 315. 57. 6.5 3 329. 52. 6.2 14 312. 64. 7.1 4 330. 55. 7.0 15 316. 57. 5.7 5 330. 52. 6.2 16 312. 60. 6.0 6 325. 48. 7.0 17 310. 60. 5.3 7 322. 47. 8.2 18 311. 58. 6.0 8 325. 55. 8.1 19 320. 44. 5.4 9 325. 52. 8.2 20 340. 48. 6.0 10 319. 62. 6.8 73/46 21 330. 45. 5.8 8 331. 59. 6.4 22 329. 47. 5.2 9 327. 68. 6.2 23 315. 65. 4.6 10 327. 59. 6.1 0 305. 54. 5.0 11 328. 72. 5.8 1 308. 49. 5.0 12 326. 67. 6.1 2 316. 62. 4.8 13 330. 61. 6.5 CPS/USAR CHAPTER 02 2.3-22 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 3 309. 56. 4.3 14 332. 62. 6.3 4 311. 54. 5.8 15 334. 57. 5.9 5 315. 54. 5.4 16 332. 55. 5.8 6 320. 37. 6.0 17 335. 36. 5.7 7 330. 46. 6.3 18 337. 39. 5.4 73/48 60M DIR. RNG. SPD DIR. RNG. SPD 3 45. 60. 3.2 10 194. 70. 7.2 4 219. 59. 2.8 11 193. 76. 7.5 5 268. 9. 2.8 12 189. 63. 7.0 6 271. 39. 4.2 13 194. 56. 7.0 7 277. 35. 3.3 14 202. 69. 7.3 8 255. 61. 2.8 15 204. 76. 6.9 9 155. 105. 4.7 16 200. 54. 7.1 10 163. 95. 5.4 17 195. 57. 7.0 11 178. 87. 6.7 18 199. 54. 5.6 12 168. 87. 6.9 19 187. 59. 5.6 13 179. 89. 7.4 20 189. 67. 7.3 14 176. 90. 6.8 21 198. 51. 6.0 15 180. 91. 6.5 22 216. 53. 5.5 16 179. 77. 6.2 23 209. 47. 5.7 17 185. 92. 6.3 0 205. 57. 4.2 18 177. 82. 6.4 1 210. 50. 4.1 19 173. 65. 6.6 2 201. 42. 5.1 20 177. 52. 6.4 3 210. 41. 5.4 21 179. 52. 6.5 4 216. 44. 5.8 22 181. 61. 7.0 5 204. 54. 6.1 23 181. 61. 6.0 6 222. 58. 5.5 0 183. 65. 6.1 7 206. 41. 5.5 1 199. 62. 6.0 8 211. 122. 5.1 2 206. 55. 7.1 9 208. 61. 5.2 3 207. 59. 7.2 10 230. 85. 5.5 4 201. 56. 6.2 11 236. 63. 6.3 5 187. 53. 5.7 12 228. 66. 7.2 6 186. 52. 6.1 13 216. 78. 8.2 7 183. 63. 6.3 14 220. 74. 7.8 8 181. 52. 6.0 15 226. 67. 9.4 9 184. 58. 5.9 73/51/7 60M DIR. RNG. SPD DIR. RNG. SPD 7 297. 67. 10.4 0 193. 65. 6.7 8 298. 76. 10.1 1 206. 73. 8.9 9 296. 56. 10.2 2 232. 62. 9.2 10 296. 66. 9.6 3 250. 37. 9.6 11 295. 62. 9.7 4 272. 54. 11.1 12 305. 78. 8.9 5 288. 44. 10.4 13 296. 79. 8.4 6 295. 55. 9.3 14 283. 72. 8.6 7 299. 62. 8.9 CPS/USAR CHAPTER 02 2.3-23 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 15 285. 79. 8.2 8 307. 68. 10.2 16 278. 61. 6.7 9 304. 67. 11.2 17 265. 43. 6.8 10 306. 53. 11.9 18 249. 41. 5.9 19 238. 43. 4.5 20 199. 35. 4.7 21 189. 46. 5.4 22 193. 44. 5.7 23 194. 64. 6.4 Hour T T 10MDP60MDP 73/61 6 7.6 3.6 3.0 7 7.8 .30 4.0 3.1 8 8.0 -.09 5.0 3.5 9 8.1 -.10 6.0 4.8 10 8.2 -.14 6.8 5.8 Hour 10M DIR. RNG. 73/66 1 208. 69.

2 198. 67.

3 202. 66.

4 211. 80.

Hour T T 10MDP60MDP 73/94 12 3.8 -0.42 1.4 1.1 Hour 10M DIR. RNG. SPD 73/112 16 250. 45. 2.0 17 251. 55. 1.8 18 249. 48. 2.2 19 252. 52. 2.1 60M DIR. RNG. SPD 260. 30. 3.8 255. 35. 3.5 257. 45. 4.1 262. 40. 4.3 73/112 10MT T 10MDP60MDPP 16 16.1 0.20 8.8 8.0 0.00 17 16.0 0.45 8.7 7.9 0.00 18 15.8 0.75 8.6 7.4 0.00 19 15.7 0.80 8.4 7.3 0.00

73/120 .T T 10MDP60MDP 21 19.0 -0.20 14.1 14.6 CPS/USAR CHAPTER 02 2.3-24 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 8 Precip 8 hr. No DATA 73/128 73/135 .T T 10MDP60MDP 11 13.8 -0.85 -4.2 -4.0 12 14.6 -0.87 -4.5 -4.0

73/144/19

thru 73/152/15 Precip 236 hr. No DATA 73/168/13

thru 73/170/12 Precip 46 hr. No DATA 73/179 Hour 60M DIR. RNG. SPD Hour DIR. RNG. SPD 8 304. 58. 9.4 3 303. 78. 3.1 9 310. 53. 9.4 4 325. 36. 3.5 10 301. 60. 9.5 5 341. 108. 2.5 11 321. 71. 9.1 6 281. 68. 2.9 12 311. 73. 9.2 7 344. 73. 3.1 13 300. 71. 6.0 8 345. 131. 2.4 14 301. 51. 10.5 9 317. 65. 6.3 15 298. 69. 9.7 10 305. 74. 5.0 16 301. 55. 5.7 11 316. 134. 4.5 17 305. 59. 9.1 12 302. 96. 6.7 18 305. 51. 8.0 13 289. 100. 7.4 19 306. 48. 5.4 14 272. 80. 5.8 20 299. 51. 4.2 15 289. 40. 7.5 21 283. 34. 3.6 16 278. 75. 6.0 22 246. 29. 3.4 17 289. 54. 7.0 23 246. 28. 3.8 18 282. 49. 5.7 0 257. 20. 4.5 19 290. 104. 3.4 1 275. 32. 3.6 20 275. 32. 3.1 2 272. 37. 2.7 73/195 Hour T T 10MDP 60MDP 9 24.0 -0.65 16.8 17.2 10 25.0 -0.70 17.0 17.4 11 26.0 -0.75 17.1 17.6 73/197 Hour T T 10MDP 60MDP 15 26.3 -0.74 8.4 8.4 16 26.2 -0.70 8.2 7.9 73/211 10M DIR. RNG. SPD DIR. RNG. SPD 18 328. 79. 0.4 8 326. 55. 2.0 19 154. 69. 3.3 9 328. 83. 1.6 CPS/USAR CHAPTER 02 2.3-25 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 20 145. 44. 3.2 10 349. 63. 2.0 21 188. 57. 3.7 11 3. 53. 1.9 22 198. 59. 0.4 12 345. 70. 3.9 23 170. 42. 1.7 13 351. 68. 4.3 0 149. 70. 0.3 14 346. 67. 5.4 1 192. 30. 0.4 15 356. 53. 6.8 2 149. 14. 0.4 16 360. 55. 7.0 3 181. 0. 0.6 17 4. 45. 3.8 4 246. 19. 0.6 18 8. 42. 4.8 5 286. 0. 0.7 19 351. 39. 4.1 6 292. 25. 0.8 20 335. 39. 4.5 7 304. 59. 0.7 21 346. 60. 4.5 73/219 Hour 60M DIR. RNG. DIR. RNG.

10 201. 79. 22 172. 68.

11 200. 87. 23 170. 60.

12 187. 72. 0 183. 68. 13 192. 76. 1 182. 74. 14 191. 94. 2 183. 70.

15 194. 76. 3 187. 71.

16 188. 60. 4 193. 62.

17 179. 71. 5 190. 53.

18 168. 62. 6 188. 56. 19 130. 50. 7 193. 54. 20 144. 52. 8 195. 77.

21 147. 55.

73/228 Hour 60M DIR. RNG. SPD DIR. RNG. SPD 18 145. 50. 2.3 1 128. 44. 2.3 19 127. 26. 2.4 2 115. 34. 2.0 20 125. 46. 1.2 3 104. 53. 2.1 21 127. 67. 1.8 4 109. 47. 2.5 22 103. 59. 2.1 5 135. 28. 2.7 23 113. 61. 2.2 6 127. 29. 2.8 0 144. 52. 2.2 7 120. 51. 2.9 73/233 Hour 60M DIR. RNG. SPD DIR. RNG. SPD 10 44. 81. 5.6 23 60. 53. 4.0 11 10. 138. 4.9 0 33. 27. 4.0 12 14. 142. 5.1 1 4. 46. 4.8 13 31. 137. 5.7 2 51. 58. 4.6 14 10. 91. 5.9 3 71. 65. 4.7 15 4. 61. 6.9 4 83. 96. 4.3 16 13. 69. 6.7 5 78. 79. 3.4 17 343. 47. 4.9 6 75. 65. 3.5 18 342. 59. 4.0 7 90. 67. 5.4 19 51. 93. 4.1 8 117. 73. 6.5 20 59. 77. 3.9 9 129. 94. 7.0 CPS/USAR CHAPTER 02 2.3-26 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 21 79. 69. 5.2 10 140. 95. 6.9 22 56. 45. 3.8 11 103. 101. 5.5 73/244 Hour 60M DIR. RNG. SPD DIR. RNG. SPD 7 125. 69. 9 175. 86.

8 157. 72. 10 182. 85.

9 171. 72. 11 179. 91.

10 172. 92. 12 176. 84.

11 186. 106. 13 178. 81.

12 173. 82. 14 172. 88. 13 182. 93. 15 176. 72. 14 176. 83. 16 175. 86.

15 166. 87. 17 177. 69.

16 164. 86. 18 159. 63.

17 148. 65. 19 142. 41.

18 148. 71. 20 130. 49. 19 150. 65. 21 128. 47. 20 146. 60. 22 136. 53.

21 153. 60. 23 131. 44.

22 154. 72. 0 180. 83.

23 159. 72. 1 138. 60.

0 164. 66. 2 140. 27. 1 166. 63. 3 150. 27. 2 145. 55. 4 136. 24.

3 142. 55. 5 147. 28.

4 138. 50. 6 159. 63.

5 157. 80. 7 149. 71.

6 153. 69. 8 146. 71. 7 145. 86. 9 161. 89. 8 167. 85. 10 167. 88.

11 175. 87. 23 158. 90.

12 217. 66. 0 145. 32.

13 200. 73. 1 161. 64. 14 169. 71. 2 199. 53. 15 179. 88. 3 205. 139.

16 186. 60. 4 209. 102.

17 184. 79. 5 204. 67.

18 179. 73. 6 192. 55.

19 174. 76. 7 179. 65. 20 197. 69. 8 210. 75. 21 182. 74. 9 218. 77.

22 200. 66. 10 223. 71.

73/265 Hour 10M SPD 21 3.0 CPS/USAR CHAPTER 02 2.3-27 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 22 3.2 23 2.9 0 2.5 1 2.0 2 1.3 3 0.8 4 1.0 5 1.0 6 0.8 73/266 Hour 10M DIR. SPD DIR. SPD 19 2.4 22 4.5 20 2.8 23 5.0 21 4.0 73/269 Hour T T 10MDP60MDP 9 22.8 -0.48 17.3 16.6 10 23.0 -0.42 17.+ 16.7 11 23.1 -0.52 17.4 16.8 12 23.3 -0.45 17.5 16.9 13 23.4 -0.38 17.5 16.9 14 23.6 -0.45 17.6 16.9 15 23.8 -0.10 17.6 17.0 16 24.2 -0.59 17.6 17.1 17 24.8 -0.85 17.6 17.3 73/318 Hour 60M DIR. RNG. SPD Hour DIR. RNG. SPD 3 243. 163. 2.0 19 194. 46. 10.6 4 200. 63. 3.2 20 191. 60. 10.4 5 186. 54. 3.5 21 195. 53. 10.5 6 153. 53. 3.2 22 192. 59. 10.7 7 154. 55. 3.0 23 197. 53. 11.1 8 149. 94. 3.4 0 192. 50. 10.6 9 129. 60. 5.2 1 196. 44. 9.8 10 158. 72. 5.4 2 206. 56. 9.6 11 162. 86. 7.0 3 198. 45. 8.8 12 155. 76. 7.1 4 194. 60. 9.7 13 163. 66. 7.6 5 207. 64. 7.3 14 170. 71. 8.2 6 257. 94. 8.1 15 169. 64. 8.5 7 180. 33. 5.1 16 176. 54. 8.8 8 213. 65. 7.5 17 182. 58. 9.8 9 261. 35. 11.5 18 188. 50. 10.1 10 262. 55. 11.7 73/319 Hour 60M DIR. RNG. SPD Hour DIR. RNG. SPD 22 311. 51. 5.5 8 269. 30. 4.0 CPS/USAR CHAPTER 02 2.3-28 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 23 285. 55. 4.5 9 298. 66. 5.1 0 262. 34. 4.6 10 293. 62. 5.0 1 291. 49. 7.4 11 289. 63. 5.0 2 311. 46. 8.1 12 283. 76. 5.0 3 316. 39. 8.3 13 284. 56. 5.1 4 308. 46. 6.0 14 264. 108. 4.3 5 298. 45. 5.0 15 247. 77. 3.5 6 300. 60. 4.0 16 210. 105. 3.0 7 251. 37. 3.0 73/320 Hour 10M SPD Hour SPD 17 0.6 19 0.4 18 0.4 20 1.5 73/320/17 - 73/324/10 Precip No DATA 73/333 Hour 60M DIR. RNG. SPD Hour DIR. RNG. SPD 11 198. 80. 2.8 13 207. 55. 6.3 12 197. 94. 4.2 14 199. 63. 5.8 73/335 60M DIR. RNG. SPD DIR. RNG. SPD 19 104. 51. 6.4 16 163. 64. 6.6 20 110. 46. 7.6 17 167. 58. 6.7 21 120. 49. 7.0 18 176. 53. 6.2 22 115. 44. 7.2 19 175. 61. 6.4 23 124. 42. 6.1 20 181. 51. 8.8 0 132. 41. 5.9 21 182. 50. 8.8 1 139. 39. 6.1 22 193. 49. 8.7 2 145. 38. 5.6 23 186. 48. 9.5 3 140. 37. 5.9 0 200. 43. 8.2 4 150. 44. 6.3 1 198. 49. 7.6 5 152. 42. 6.6 2 189. 53. 7.4 6 153. 38. 5.5 3 189. 50. 6.9 7 152. 38. 5.0 4 185. 40. 7.0 8 162. 56. 5.7 5 183. 36. 7.1 9 168. 60. 6.6 6 182. 43. 7.2 10 182. 50. 7.3 7 185. 38. 7.2 11 176. 73. 7.5 8 186. 50. 6.4 12 177. 62. 8.0 9 190. 43. 7.3 13 177. 63. 9.5 10 195. 54. 8.1 14 175. 66. 9.1 11 200. 48. 8.3 15 167. 64. 8.5 73/346 Hour 60M SPD Hour SPD 1 5.8 6 2.2 2 4.7 7 2.5 3 2.2 8 1.5 4 3.5 9 3.1 5 3.3 10 3.0 CPS/USAR CHAPTER 02 2.3-29 REV. 11, JANUARY 2005 ATTACHMENT B (CONT'D)

(Q&R 451.02)

Date Hour Parameter Hour Parameter 73/358 Hour 10M DIR. RNG Hour DIR. RNG. 21 116. 37. 8 172. 30. 22 116. 50. 9 157. 44. 23 119. 51. 10 176. 50. 0 127. 39. 11 185. 42.

1 136. 47. 12 214. 57.

2 149. 47.

3 151. 48.

4 175. 26.

5 181. 40. 6 174. 25. 7 172. 27.

73/361 Hour 10M SPD Hour SPD 3 3.4 7 3.4 4 3.6 8 3.5 5 3.5 9 4.0 6 3.8 74/10 Hour 10M SPD Hour SPD 19 3.6 3 3.4 20 3.2 4 4.1 21 3.2 5 3.8 22 2.6 6 3.7 23 3.0 7 2.0 0 3.1 8 2.2 1 3.7 9 3.0 2 3.6 CPS/USAR CHAPTER 02 2.3-30 REV. 11, JANUARY 2005 2.3.4 Short-Term Diffusion Estimates 2.3.4.1 Objective Conservative estimates of the local atmospheric dilution factors (X/Q) and their 5% probability level conditions for the Clinton Power Station site have been prepared for the exclusion area boundary (EAB), actual site boundary (ASB), and distances of 0.5, 1.5, 2.5, 3.5, 4.5, 7.5, 15, 25, 35, and 45 miles. Calculations were made for sliding time period windows of 1, 8, 16, 72, and 624 hours0.00722 days <br />0.173 hours <br />0.00103 weeks <br />2.37432e-4 months <br /> from onsite meteorological data for the period May 1972 through April 1977. 2.3.4.2 Calculations Calculations of ground-level atmospheric dilution factors for the CPS site were performed using Gaussian plume diffusion models for a continuously emitting ground level source. Hourly centerline X/Q values were computed from the concurrent hourly mean values of wind speed, wind direction and range, and Pasquill stability class of the onsite meteorological data. The wind speed at the 10 meter level was used in the diffusion estimates for the ground-level release. The Pasquill stability class was determ ined from the measured vertical temperature difference (T) and the variance of the horizontal wind field () according to Regulatory Guide 1.23. Calms were assigned a wind speed value equal to the starting speed of the wind vane (0.7 mph). Cumulative frequency distributions were prepared to determine the /Q values that were exceeded 5% and 50% of the time. 2.3.4.3 Atmospheric Diffusion Models and Frequency Distributions Gaussian plume diffusion models for ground-level concentration were used to describe the downwind spread of effluents for the Clinton Power Station. A continuous ground-level release of effluents at a constant emission rate was assumed in the diffusion estimates. Total reflection of the plume at ground-level was assumed in the diffusion estimates: i.e., there is no deposition or reaction at the surface. Hourly X/Q values were calculated by the following equations: zy10u1QX= (2.3-1) )2/A(u1QXzy10+= (2.3-2) or )3(u1QXzy10= (2.3-3) where CPS/USAR CHAPTER 02 2.3-31 REV. 12, JANUARY 2007 QX is the relative centerline concentration (sec/m

3) at ground level is 3.14159 10u is the wind speed (m/sec) at 10 meters above the ground y is the lateral plume spread (m), a function of atmospheric stability, wind speed, and downwind distance from the point of release. For distances to 800 meters,yyM=; M being a function of atmospheric stability and wind speed. For distances greater than 800 meters, Sy = (M-1) sy800m + sy y is the lateral plume spread as a function of atmospheric stability and distance z is the vertical plume spread as a function of atmospheric stability and distance. is the smallest vertical plane, cross-sectional area of the building from which the effluent is released (A=2069m 2). For neutral to stable conditions with wind speeds less than 6m/sec Equations 2.3-2 and 2.3-3 were calculated and compared, and the higher X/Q was selected. This higher value was compared to the X/Q resulting from Equation 2.3-1 and the lower was selected. This was done in accordance with Regulatory Guide 1.145, Atmospheric Dispersion Models For Potential Accident Consequence Assessments At Nuclear Power Plants. For all other stability and/or wind speed conditions, X/Q was selected as the higher value from Equations 2.3-2 and 2.3-3. From these hourly X/Q values, cumulative frequency distributions were prepared from the mean values of sliding time windows of 1, 2, 8, 16, 72, and 624 hours0.00722 days <br />0.173 hours <br />0.00103 weeks <br />2.37432e-4 months <br />. These intervals correspond to time periods of 0-1 hour, 0-2 hours, 0-8 hours, 8-24 hours, 1-4 daysand 4-30 days. For each time period used, the mean centerline X/Q value in each sector was computed. The results are presented in Tables 2.3-30 through 2.3-43.

2.3.5 Short-Term (Accident) Diffusion Estimates (Alternative Source Term X/Q Analysis) 2.3.5.1 Objective Estimates of atmospheric diffusion (X/Q) at the Exclusion Area Boundary (EAB), the outer boundary of the Low Population Zone (LPZ) and the Control Room Intakes have been prepared for the regulated short-term (accident) time-averaging periods of 0-2 hrs, 2-8 hrs, 8-24 hrs, 1-4 days and 4-30 days. Calculations were made based on onsite meteorological data for the years 2000 through 2002. 2.3.5.2 Calculation of X/Q at the EAB and LPZ X/Q was calculated at the EAB (975 m) and LPZ (4018 m) for the Standby Gas Treatment/HVAC Vent Stack using the NRC-recommended model PAVAN (Reference 27).

CPS/USAR CHAPTER 02 2.3-31a REV. 12, JANUARY 2007 This stack does not qualify as an elevated release as defined by Regulatory Guide 1.145 (Reference 28); therefore, it is executed as a "ground" type release. X/Q values at the EAB and LPZ were calculated by PAVAN in accordance with Regulatory Guide 1.145. For ground-level releases, calculation for the 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> following the accident is based on the following equations:

()2110AUQzy+= (2.3.5-1)

()zyUQ3110= (2.3.5-2) zyUQ=101 (2.3.5-3) where: Q/ is relative concentration, in sec/m

3. is 3.14159.

10U is wind speed at 10 meters above plant grade, in m/sec.

y is lateral plume spread, in meters, a function of atmospheric stability and distance.

z is vertical plume spread, in meters, a function of atmospheric stability and distance.

y is lateral plume spread with meander and building wake effects (in meters), a function of atmospheric stability, wind speed, and distance [for distances of 800 m or less, y= My, where M is determined from Reg. Guide 1.145 Fig. 3; for distances greater than 800 m, y=(M-1) y800 m +y. A is the smallest vertical-plane cross-sectional area of the reactor building, in m

2. (Other structures or a directional consideration may be justified when appropriate.) Plume meander is only considered during neutral (D) or stable (E, F, or G) atmospheric stability conditions. For such, the higher of the values resulting from Equations 2.3.5-1 and 2.3.5-2 is compared to the value of Equation 2.3.5-3 for meander, and the lower value is selected. For all other conditions (stability classes A, B, or C), meander is not considered and the highest X/Q value of equations 2.3.5-1 and 2.3.5-2 is selected. The X/Q values calculated at the EAB based on meteorological data representing a 1-hour average are assumed to apply for the entire 2-hour period.

CPS/USAR CHAPTER 02 2.3-31b REV. 12, JANUARY 2007 To determine the "maximum sector 0-2 hour X/Q" value at the EAB, PAVAN constructs a cumulative frequency probability distribution (probabilities of a given X/Q value being exceeded in that sector during the total time) for each of the 16 sectors using the X/Q values calculated for each hour of data. This probability is then plotted versus the X/Q values and a smooth curve is fitted to form an upper bound of the computed points. For each of the 16 curves, the X/Q value that is exceeded 0.5 percent of the total hours is selected and designated as the sector X/Q value. The highest of the 16 sector X/Q values is the maximum sector X/Q. Determination by PAVAN of the LPZ maximum sector X/Q is based on a logarithmic interpolation between the 2-hour sector X/Q and the annual average X/Q for the same sector. For each time period, the highest of these 16 sector X/Q values is identified as the maximum sector X/Q value. The maximum sector X/Q values will, in most cases, occur in the same sector. If they do not occur in the same sector, all 16 sets of values are used in dose assessment requiring time-integrated concentration considerations. The set that results in the highest time-integrated dose within a sector is considered the maximum sector X/Q. The "5% overall site X/Q" values for the EAB a nd LPZ are each determined by constructing an overall cumulative probability distribution for all directions. The 0-2 hour X/Q values computed by PAVAN are plotted versus their probability of being exceeded, and an upper bound curve is fitted by the model. From this curve, the 2-hour X/Q value that is exceeded 5% of the time is determined. PAVAN then calculates the 5% overall site X/Q at the LPZ for intermediate time periods by logarithmic interpolation of the maximum of the 16 annual average X/Q values and the 5% 2-hour X/Q values. 2.3.5.2.1 PAVAN Meteorological Databases The meteorological database to be utilized for the EAB and LPZ X/Q calculations were prepared for use in PAVAN by transforming the three years (i.e. 2000-2002) of hourly meteorological tower data observations into a joint wind speed-wind direction-stability class occurrence frequency distribution as shown in Tables 2.3-45 and 2.3-46. In accordance with Regulatory Guide 1.145, atmospheric stability class was determined by vertical temperature difference between the 60 m and the 10-m level, and wind direction was distributed into 16- 22.5° sectors. Seven (7) wind speed categories were defined according to Regulatory Guide 1.23 (Reference 29) with the first category identified as "calm". The higher of the starting speeds of the wind vane and anemometer (i.e. 0.50 mph) was used as the threshold for calm winds, per Regulatory Guide 1.145, Section 1.1. A midpoint was also assumed between each of the Regulatory Guide 1.23 wind speed categories, Nos. 2-6, as to be inclusive of all wind speeds. The wind speed categories have therefore been defined as follows: PAVAN WIND SPEED CATEGORIES Category No.

Regulatory Guide 1.23 Speed Interval (mph)

Pavan-Assumed Speed Interval (mph) 1 (Calm) 0 to < 1 0 to <0.50 2 1 to 3 >=0.50 to <3.5 3 4 to 7 >=3.5 to <7.5 4 8 to 12 >=7.5 to <12.5 CPS/USAR CHAPTER 02 2.3-31c REV. 12, JANUARY 2007 5 13 to 18

>=12.5 to <18.5 6 19 to 24

>=18.5 to <24 7 >24 >=24 The procedures used by PAVAN assign a direction to each calm hour according to the directional distribution for the lowest non-calm wind-speed class. This procedure is performed separately for the calms in each stability class. 2.3.5.2.2 PAVAN Model Input Parameters The Standby Gas Treatment/HVAC Vent Stack has height of 60.5 m above station grade, however since it does not qualify as an elevated release per Regulatory Guide 1.145, PAVAN requires that its height be assigned an input value of 10 m. For this assumed non-elevated stack scenario, EAB and LPZ receptor terrain elevation is not considered. The smallest Control Building vertical projected area of the 1093.5 m 2 was utilized (h=33.5 m, w=32.6 m; based on drawing M01-1115. 2.3.5.2.3 PAVAN EAB and LPZ X/Q Modeling Results Atmospheric X/Q diffusion estimates predicted by PAVAN at the EAB and LPZ are summarized below. OFFSITE X/Q SUMMARY (sec/m

3) Standby Gas Treatment Vent / HVAC Vent Stack Receptor 0-2 hour 2-8 hour 8-24 hour 1-4 day 4-30 day EAB (975 m) 2.46E-04 1.19E-04 8.30E-05 3.78E-05 1.22E-05 LPZ (4018 m) 5.62E-05 2.48E-05 1.65E-05 6.81E-06 1.91E-06 2.3.5.3 Calculation of X/Q at the Control Room Intake Estimates of atmospheric diffusion (X/Q) are made at each of the three Control Room Intakes (i.e. East, West, and Normal) for releases from the Standby Gas Treatment Vent/HVAC Vent Stack for periods of 2, 8, and 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> and for 3 and 26 days. The NRC-sponsored computer code ARCON96 (Reference 30), consistent with the procedures in Draft Regulatory Guide DG-1111 (Reference 31) is utilized. 2.3.5.3.1 ARCON96 Model Analysis Since the Standby Gas Treatment/HVAC Vent Stack is not 2.5 times the height of the adjacent structures, it does not qualify as an elevated release per DG-1111, therefore, ARCON96 is executed in vent release mode. With an assumed zero (0) vertical exit velocity, vent releases are treated as ground-level releases by ARCON96. The basic model for a ground-level release is CPS/USAR CHAPTER 02 2.3-31d REV. 12, JANUARY 2007 =2yzyy0.5expU1Q (2.3.5-4) where: X/Q = relative concentration (concentration divided by release rate) [{ci/m 3)/(ci/s)] y, z = diffusion coefficients (m)

U = wind speed (m/s) y = distance from the center of the plume (m) This equation assumes that the release is continuous, constant, and of sufficient duration to establish a representative mean concentration. It also assumes that the material being released is reflected by the ground. Diffusion coefficients are typically determined from atmospheric stability and distance from the release point using empirical relationships. A diffusion coefficient parameterization from the NRC PAVAN and XOQDOQ (Reference 32) codes is used for y and z. The diffusion coefficients have the general form

= a x b + c were x is the distance from the release point, in meters, and a , b, and c are parameters that are functions of stability. The parameters are defined for 3 distance ranges - 0 to 100 m, 100 to 1000 m, and greater than 1000 m. The parameter values may be found in the listing of Subroutine NSIGMA1 in Appendix A of NUREG/CR-6331 Rev. 1. Diffusion coefficient adjustments for wakes and low wind speeds are incorporated as follows: To estimate diffusion in building wakes, composite wake diffusion coefficients, y and z, replace y and z. The composite wake diffusion coefficients are defined by 1/22y22y12yy++= (2.3.5-5) 1/22z22z12zz++= (2.3.5-6)

The variables y and z are the normal diffusion coefficients, y1 and z1 are the low wind speed corrections, and y2 and z2 are the building wake corrections. These corrections are described and evaluated in Ramsdall and Fosmire (Reference 33). The low wind speed corrections are:

CPS/USAR CHAPTER 02 2.3-31e REV. 12, JANUARY 2007 +x=1000Ux-exp1000Ux11109.1352y1 (2.3.5-7) +x=100Ux-exp100Ux111067.622z1 (2.3.5-8) The variable x is the distance from the release point to the receptor, in meters, and U is the wind speed in meters per second. It is appropriate to use the slant range distance for x because these corrections are made only when the release is assumed to be at the ground level and the receptor is assumed to be on the axis of the plume. The diffusion coefficients corrections that account for enhanced diffusion in the wake have a similar form. These corrections are: +x=A10x-expA10x11AU1024.522-2y2 (2.3.5-9) +x=A10x-expA10x11AU1017.122-2z2 (2.3.5-10) The constant A is the cross-sectional area of the building. An upper limit is placed on y as a conservative measure. This limit is the standard deviation associated with a concentration uniformly distributed across a sector with width equal to the circumference of a circle with radius to the distance between the source and receptor. This value is 12x2ymax= x81.1 (2.3.5-11) 2.3.5.3.1.1 ARCON96 Meteorological Databases The 2000-2002 meteorological databases utilized in ARCON96 consists of hourly meteorological data observations of wind speed and direction, and delta temperature stability class. The designation of 'calm' is made to all wind speed observations of less than 0.5 mph. The higher of the starting speeds of the Climatronics wind vane and anemometer equipment on each of the towers (i.e. 0.50 mph) was used as the threshold for calm winds, per Regulatory Guide 1.145, Section 1.1. 2.3.5.3.1.2 ARCON96 Input Parameters The parameters that were input into the ARCON96 model for use in calculating the Control Room X/Q are summarized below:

CPS/USAR CHAPTER 02 2.3-31f REV. 12, JANUARY 2007 ARCON96 MODEL INPUT PARAMETERS ARCON96 INPUT PARAMETER East Intake West Intake Normal Intake Release Height (m) 60.5 60.5 60.5 Intake Height (m) 29.9 18.6 28.3 Horizontal Distance from Intake to Stack (m) 69.8 69.8 51.5 Elevation Difference between Stack Grade and Intake Grade (m) 0 0 0 Building Area (m

2) 1093.5 1093.5 1093.5 Direction from Intake To Stack (°)

288 168 260 Vertical Velocity (m/s) 0 0 0 Stack Flow (m 3/s) 0 0 0 Stack Radius (m) 0 0 0 2.3.5.3.1.3 ARCON96 Control Room X/Q Results A summary of the atmospheric diffusion estimates at the Control Room Intakes for releases from the Standby Gas Treatment/HVAC Vent Stack is shown below. ARCON96 Control Room Intake X/Q Results (sec/m

3) Standby Gas Treatment/HVAC Vent Stack INTAKE 0-2 Hour 2-8 Hour 8-24 Hour 1-4 Day 4-30 Day East Intake 9.75E-04 7.09E-04 2.93E-04 2.13E-04 1.79E-04 West Intake 9.45E-04 7.58E-04 3.28E-04 2.61E-04 1.85E-04 Normal Intake

• 1.54E-03 1.09E-03 4.67E-04 3.21E-04 2.64E-04 * (maximum intake X/Q value)

CPS/USAR CHAPTER 02 2.3-31g REV. 12, JANUARY 2007 2.3.6 Long-Term (Routine) Diffusion Estimates 2.3.6.1 Objective Annual average dilution factors were computed for routine releases from the common station vent along the side of the containment building. The MESODIF model was used. Meteorological data observed on the tower at the Clinton site were used. The period of record was May 14, 1972 through April 30, 1977. 2.3.6.2 Calculations MESODIF employs an integrated puff model concept. This model differs from ordinary Gaussian type models in that it will allow released materials to be transported back over the source in the event of a wind shift. MESODIF carries the effluent as a string of puffs released into the wind field as observed by the onsite meteorological station. Individual puffs are tracked until they are either too dilute to be of further significance or else leave the area being CPS/USAR CHAPTER 02 2.3-32 REV. 11, JANUARY 2005 considered. The integrated puff concept yields a conservative estimate of concentration near the source. Ground level releases were assumed in order to yield conservative estimates. The MESODIF program developed by the Air Resources Laboratories (ARL) personnel at Idaho Falls, Idaho was described by Start and Wendell (Reference 26). A program source deck was obtained from ARL in January 1978. Modifications were made to the program to accommodate input data from a single site rather than a number of stations as used in Idaho. The modifications are described in the following paragraphs. Subroutines RNGRD9 and ASCND were deleted from the program. Subroutine ASCND was used to move elements of an array. Subroutine RNGRD9 was used to read the wind direction and speed data, convert the direction and speed to U- and V-components and to interpolate the components from station locations to a grid array. Meteorological data for stability and mixing depth were read in the main program. The array in the main program has space available for wind direction and speed so these data were supplied there and conversion to U- and V-components was accomplished in the main program. Wind direction was provided to the nearest degree and wind speed to the nearest mile per hour.

Conversion to U- and V-components was accomplished by: = (270 - WD) /180 U = S cos V = S sin where WD is wind direction and S is wind speed for any hour. The U- and V-components calculated in this manner were assigned to each grid point. Two U, V arrays are carried in the program because an interpolation is performed to account for changes with time. This modification maintains both arrays at two times. A test case supplied by ARL was run before and after the change. Constant wind direction and speeds were assumed at all stations for the "before" run. Identical results were achieved for the

test runs. Hourly data from May 11, 1972 through April 30, 1977 were used. Integrated dosages were calculated for each year and the hourly values averaged for the five year period. The rectangular array of points from 2 mile and 10 mile grids were plotted. Sector centerline values were derived from the data and plotted on log-log graph paper. A straight line was drawn through the points and the relative concentrations were read at the required distances. Actual model calculations were made at distances ranging from two miles to 45 miles from the source. Data are listed in Table 2.3-44 for the period of record.

CPS/USAR CHAPTER 02 2.3-33 REV. 12, JANUARY 2007

2.3.7 References

1. L. Denmark, "Climates of the States: Illinois," No. 60-11, U.S. Department of Commerce, August 1959 (Revised June 1969). 2. "Local Climatological Data, Annual Summary with Comparative Data, Peoria, Illinois, 1976," U.S. Department of Commerce, NOAA, Asheville, North Carolina. 3. "Local Climatological Data, Annual Summary with Comparative Data, Springfield, Illinois, 1976," U.S. Department of Commerce, NOAA, Asheville, North Carolina. 4. R. A. Bryson, "Airmasses, Streamlines and the Boreal Forest," Technical Report No. 24, pp. 13-57, University of Wisconsin: Dept. of Meteorology, Madison, Wisconsin, 1966. 5. S. A. Changnon, Jr., "Climatology of Hourly Occurrences of Selected Atmospheric Phenomena in Illinois," Circular 93, Illinois State Water Survey, Urbana, Illinois, 1968. 6. S. A. Changnon, Jr., Thunderstorm-Precipitation Relations in Illinois, Report of Investigation No. 34, Illinois State Water Survey, Urbana, Illinois, 1957. 7. Glossary of Meteorology, (Edited by R. E. Huschke), Second Printing with Corrections, American Meteorological Society, Boston, Massachusetts, 1970. 8. "Severe Local Storm Occurrences, 1955-1967," WBTM FCST 12, U.S. Department of Commerce, ESSA, Silver Spring, Maryland, September 1969. 9. F. A. Huff and S. A. Changnon, Jr., Hail Climatology of Illinois, Report of Investigation 38, Illinois State Water Survey, Urbana, Illinois, 1959. 10. J. L. Marshall, "Probability of a Lightning Stroke," Lightning Protection, Chapter 3, pp. 30-31, John Wiley and Sons, New York, New York, 1971. 11. J. W. Wilson and S. A. Changnon, Jr., "Illinois Tornadoes," Circular 103, Illinois State Water Survey, Urbana, Illinois, 1971. 12. J. C. S. Thom, "Tornado Probabilities," Monthly Weather Review, Volume 91, pp. 730-736, 1963. 13. "Design Basis Tornado for Nuclear Power Plants," Regulatory Guide 1.76, U.S. Atomic Energy Commission, April 1974. 14. "Building Code Requirements for Minimum Design Loads in Buildings and Other Structures," ANSI A58.1-1972, American National Standards Institute, Inc., New York, New York, 1972. 15. S. A. Changnon, Jr., "Climatology of Severe Winter Storms in Illinois," Bulletin 53, Illinois State Water Survey, Urbana, Illinois, 1969. 16. "Glaze Storm Loading Summary, 1927-28 to 1936-37," Association of American Railroads, 1955.

CPS/USAR CHAPTER 02 2.3-34 REV. 12, JANUARY 2007 17. "Snow Load Studies," Paper 19, Housing and Home Finance Agency, Division of Housing Research, and U.S. Weather Bureau, Washington, D.C., 1952. 18. J. T. Riedel, J. F. Appleby, and R. W. Schloemer, "Seasonal Variation of the Probable Maximum Precipitation East of the 105th Meridian for Areas from 10 to 1000 Square Miles and Durations of 6, 12, 24, and 48 Hours," HMR No. 33, U.S. Department of Commerce, Washington, D.C., April 1956. 19. C. R. Hosler, "Low-Level Inversion Frequency in the Contiguous United States," Monthly Weather Review, Volume 89, pp. 319-339, September 1961. 20. G. C. Holzworth, "Mixing Heights, Wind Speeds, and Potential for Urban Air Pollution Throughout the Contiguous United States," AP-101, U.S. Environmental Protection Agency, Office of Air Programs, Research Triangle Park, North Carolina, January 1972. 21. R. E. Huschke (Ed.), "Glossary of Meteorology," American Meteorological Society, Boston, Massachusetts, 1959 (Second Printing with Corrections, 1970). 22. U.S. Department of Commerce, NOAA, Environmental Data Service, "Local Climatological Data, Annual Summary with Comparative Data," Springfield and Peoria, Illinois (two documents). 23. Texas Water Development Board, "Simulation of Water Quality in Streams and Canals,"

Report 128, May 1971. 24. "Prediction of Thermal Energy Distribution in Streams and Reservoirs," Water Resources Engineers, Inc., Prepared for the Department of Fish and Game, State of California, August 30, 1968. 25. G. E. Harbeck, et al., "Water-Loss Investigations: Lake Mead Studies," Geological Survey Professional Paper 298, U.S. Govt. Printing Office, Washington, D.C., 1958. 26. G. E. Start and L. L. Wendell, "Regional Effluent Dispersion Calculations Considering Spatial and Temporal Meteorological Variations," NDAA Technical Memorandum ERL-ARL-44, Air Resources Laboratories, Idaho Falls, Idaho, 1974. 27. Atmospheric Dispersion Code System for Evaluating Accidental Radioactivity Releases from Nuclear Power Stations; PAVAN, Version 2; Oak Ridge National Laboratory; U.S. Nuclear Regulatory Commission; December 1997. 28. Regulatory Guide 1.145; Atmospheric Dispersion Models for Potential Accident Consequence Assessments at Nuclear Power Plants (Revison 1); U.S. Nuclear Regulatory Commission; November 1982. 29. Regulatory Guide 1.23 (Safety Guide 23), Onsite Meteorological Programs; U. S. Nuclear Regulatory Commission; USNRC Office of Standards Development; Washington, D.C.; 1972. 30. Atmospheric Relative Concentrations in Building Wakes; NUREG/CR-6331, PNNL-10521, Rev. 1; prepared by J. V. Ramsdell, Jr., C. A. Simmons, Pacific Northwest National Laboratory; prepared for U.S. Nuclear Regulatory Commission; May 1997 (Errata, July 1997).

CPS/USAR CHAPTER 02 2.3-34a REV. 12, JANUARY 2007 31. Draft Regulatory Guide DG-1111; Atmospheric Relative Concentrations for Control Room Radiological Habitability Assessments at Nuclear Power Plants; U.S. Nuclear Regulatory Commission; December 2001. 32. XOQDOQ: Computer Program for the Meteorological Evaluation of Routine Releases at Nuclear Power Stations; NUREG/CR-2919; J. F. Sagendorf, J. T. Goll, and W. F. Sandusky, U.S. Nuclear Regulatory Commission; Washington, D.C; 1982. 33. Atmospheric Dispersion Estimates in the Vicinity of Buildings; J. V. Ramsdell and C. J. Fosmire, Pacific Northwest Laboratory; 1995.

CPS/USAR CHAPTER 02 2.3-35 REV. 11, JANUARY 2005 TABLE 2.3-1 CLIMATOLOGICAL DATA FROM WEATHER STATIONS SURROUNDING CLINTON POWER STATION PARAMETER STATION PEORIA SPRINGFIELD Temperature (

°F) Annual average 50.8 52.7 Maximum 103 (July 1940) 112 (July 1954)

Minimum -20 (Jan. 1963) -22 (Feb. 1963) Degree days 6098 5558 Relative Humidity (%) Annual average at: 6 a.m. 83 82 12 noon 62 60 Wind Annual average speed (mph) 10.3 11.4 Prevailing Direction S S Fastest mile:

Speed (mph) 75 (July 75 (June Direction NW 1953) SW 1957)

Precipitation (in.) Annual average 35.06 35.02 Monthly maximum 13.09 (Sept. 1961) 9.91 (Apr. 1964)

Monthly minimum 0.03 (Oct. 1964) 0.15 (Dec. 1955) 24-hour maximum 5.06 (Apr. 1950 5.12 (Sept. 1959)

Snowfall (in.) Annual average 23.4 22.3 Monthly maximum 18.9 (Dec. 1973) 22.7 (Dec. 1973)

Maximum 24-hour 10.2 (Dec. 1973) 10.9 (Dec. 1973) Mean Annual (no. of days) Precipitation 0.1 in 111 112 Snow, sleet, hail 1.0 in.

8 8 Thunderstorms 49 50 Heavy fog (visibility 1/4mile or less) 21 18 Maximum temperature 90° F 17 28 Minimum temperature 32° F 132 119

• The data presented in this table are based upon References 2 and 3. These statistics are based on periods of record ranging from 17 to 39 years in length. The ranges span the years 1937 to 1976

.

CPS/USAR CHAPTER 02 2.3-36 REV. 11, JANUARY 2005 TABLE 2.3-2 MEASURES OF GLAZING IN VARIOUS SEVERE WINTER STORMS FOR THE STATE OF ILLINOIS STORM DATE RADIAL THICKNESS OF ICE ON WIRE (in.) RATIO OF ICE WEIGHT TO WEIGHT OF 0.25-in.

TWIG WEIGHT OF ICE (oz.) ON 1 FOOT OF STANDARD (No. 12)

WIRE CITY STATE SECTION 2-4 Feb. 1883 11 Springfield WSW 20 Mar. 1912 0.5 Decatur C 21 Feb. 1913 2.0 La Salle NE 12 Mar. 1923 1.6 12 Marengo NE 17-19 Dec. 1924 1.2 15:1 8 Springfield WSW 22-23 Jan. 1927 1.1 2 Cairo SE 31 Mar. 1929 0.5 Moline NW 7-8 Jan. 1930 1.2 Carlinville WSW 1-2 Mar. 1932 0.5 Galena NW 7-8 Jan. 1937 1.5 Quincy W 31 Dec. 1947 - 1 Jan. 1948 1.0 72 Chicago NE 10 Jan. 1949 0.8 Macomb W 8 Dec. 1956 Alton WSW 20-22 Jan. 1959 0.7 12:1 Urbana E 26-27 Jan. 1967 1.7 17:1 40 Urbana E NOTE: Based on Reference 15.

CPS/USAR CHAPTER 02 2.3-37 REV. 11, JANUARY 2005 TABLE 2.3-3 WIND-GLAZE THICKNESS RELATIONS FOR FIVE PERIODS OF GREATEST SPEED AND GREATEST THICKNESS FIVE PERIODS WHEN FIVE FIVE PERIODS WHEN FIVE FASTEST 5-MINUTE SPEEDS GREATEST ICE THICKNESSES WERE REGISTERED WERE MEASURED SPEED ICE THICKNESS ICE THICKNESS SPEED RANK (mph) (in.) (in.) (mph) 1 50 0.19 2.87 30 2 46 0.79 1.71 18 3 45 0.26 1.50 21 4 40 0.30 1.10 28 5 35 0.78 1.00 18 NOTE: From data collected throughout the United States during the period 1926-1937. Based on Reference 15.

CPS/USAR CHAPTER 02 2.3-38 REV. 11, JANUARY 2005 TABLE 2.3-4 10M TEMPERATURE (DEG. C)

AVERAGE DAILY AVERAGE DAILY MAXIMUM AVERAGE DAILY MINIMUM ABSOLUTE MAXIMUM ABSOLUTE MINIMUM January -5.1 -1.3 -8.9 15.5 -28.8 February -1.3 1.9 -4.4 15.8 -23.6 March 5.9 10.5 1.6 25.5 -15.1 April 11.4 16.7 6.1 29.3 -6.5 May 16.4 21.2 11.2 32.1 0.0 June 21.2 26.1 16.0 33.0 5.0 July 23.6 28.4 18.5 35.2 8.1 August 22.1 26.8 17.4 23.2 9.1 September 17.7 22.8 12.7 33.3 0.8 October 11.9 17.1 6.9 30.0 -4.8 November 4.5 8.4 0.8 23.0 -15.8 December -2.3 1.3 -5.9 17.8 -23.8 Period of Record 10.5 15.0 6.0 35.2 -28.8 CPS/USAR CHAPTER 02 2.3-39 REV. 11, JANUARY 2005 TABLE 2.3-5 HOURS WITH xC TEMPERATURE 32.2 DEG. OR MORE 0.0 DEG. OR LESS

-12.2 DEG. OR LESS.

-17.8 DEG. OR LESS hr. % hr. % hr % hr. %

January (0) 0.0 (2628) 72.5 (730) 20.1 (225) 6.2 February (0) 0.0 (2019) 60.5 (203) 6.1 (48) 1.4 March (0) 0.0 (808) 21.9 (19) 0.5 (0) 0.0 April (0) 0.0 (188) 4.7 (0) 0.0 (0) 0.0 May (0) 0.0 (1) 0.0 (0) 0.0 (0) 0.0 June (8) 0.2 (0) 0.0 (0) 0.0 (0) 0.0 July (67) 1.9 (0) 0.0 (0) 0.0 (0) 0.0 August (0) 0.0 (0) 0.0 (0) 0.0 (0) 0.0 September (3) 0.1 (0) 0.0 (0) 0.0 (0) 0.0 October (0) 0.0 (82) 2.3 (0) 0.0 (0) 0.0 November (0) 0.0 (948) 26.4 (28) 0.8 (0) 0.0 December (0) 0.0 (2414) 65.9 (302) 8.2 (56) 1.5 Period of Record (78) 0.2 (9088) 21.0 (1282) 3.0 (329) 0.8 CPS/USAR CHAPTER 02 2.3-40 REV. 11, JANUARY 2005 TABLE 2.3-6 DAYS WITH xC TEMPERATURE 32.2 0.0 -12.2 -17.8 DEG. OR MORE DEG. OR LESS DEG. OR LESS DEG. OR LESS DAYS % DAYS % DAYS % DAYS % January (0) 0.0 (132) 86.3 (55) 35.9 (24) 15.7 February (0) 0.0 (116) 82.3 (21) 14.9 (6) 4.3 March (0) 0.0 (65) 41.9 (2) 1.3 (0) 0.0 April (0) 0.0 (27) 16.2 (0) 0.0 (0) 0.0 May (0) 0.0 (1) 0.6 (0) 0.0 (0) 0.0 June (3) 2.0 (0) 0.0 (0) 0.0 (0) 0.0 July (15) 10.0 (0) 0.0 (0) 0.0 (0) 0.0 August (0) 0.0 (0) 0.0 (0) 0.0 (0) 0.0 September (1) 0.7 (0) 0.0 (0) 0.0 (0) 0.0 October (0) 0.0 (15) 9.9 (0) 0.0 (0) 0.0 November (0) 0.0 (73) 48.7 (3) 2.0 (0) 0.0 December (0) 0.0 (129) 83.8 (29) 18.8 (8) 5.2 Period of Record (19) 1.0 (558) 30.5 (110) 6.0 (38) 2.1 CPS/USAR CHAPTER 02 2.3-41 REV. 11, JANUARY 2005 TABLE 2.3-7 CLINTON POWER STATION SITE RELATIVE HUMIDITY SUMMARY COMPOSITE JAN FEB MAR APR MAY JUN JUL AUG SEP OCT NOV DEC PERIOD OF RECORD Average 85.94 82.04 77.29 68.01 64.44 68.24 70.00 74.04 72.15 67.15 77.58 85.71 68.28 Average Daily Max. 92.10 89.77 87.75 83.96 80.77 83.26 85.13 86.04 85.33 80.75 86.61 90.47 79.01 Average Daily Min. 71.04 65.71 56.91 46.43 43.89 47.52 49.03 53.84 49.40 45.57 60.44 71.64 50.63 Absolute Max. 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 100.00 Absolute Min. 38.34 14.11 22.26 16.80 15.78 19.22 27.20 23.93 15.91 14.86 23.13 21.40 14.11 Average for Hours 00 83.15 80.78 74.30 69.75 68.25 70.72 69.96 76.71 73.91 67.56 76.45 82.07 68.35 03 84.00 81.27 75.53 74.31 73.88 75.17 75.54 80.02 78.10 71.51 78.10 82.49 71.15 06 84.88 82.23 79.17 77.55 75.88 76.23 77.75 82.62 80.27 74.87 79.87 83.10 73.04 09 84.31 79.85 71.60 66.35 61.19 64.77 66.22 73.67 73.38 68.40 77.39 82.10 66.35 12 78.10 75.28 63.31 54.95 52.41 53.97 55.67 61.81 59.77 56.74 67.48 77.51 57.85 15 74.32 71.11 59.83 53.07 49.43 50.32 50.25 56.39 51.12 49.93 63.62 74.12 53.79 18 78.53 75.99 64.18 54.48 52.14 52.18 54.35 61.51 56.89 53.79 69.04 79.07 57.52 21 81.66 78.76 63.76 63.76 61.91 61.11 65.27 70.98 67.38 62.08 74.42 81.32 64.26 CPS/USAR CHAPTER 02 2.3-42 REV. 11, JANUARY 2005 TABLE 2.3-8 CLINTON POWER STATION SITE WET BULB TEMPERATURE SUMMARY

°C COMPOSITE JAN FEB MAR APR MAY JUN JUL AUG SEP OCT NOV DEC PERIOD OF RECORD Average -4.13 0.43 8.67 13.7418.84 23.1925.4923.64 19.9514.10 6.23 -1.22 11.34 Average Daily Max. 0.27 4.31 13.36 19.2123.52 27.3229.2827.61 24.7919.38 10.42 2.65 15.41 Average Daily Min. -7.98 -3.16 3.60 7.16 12.27 16.6919.1017.67 13.757.95 1.80 -5.00 6.33 Absolute Max. 16.67 17.76 27.70 32.1333.00 34.1735.5932.41 33.1533.13 24.67 18.67 35.59 Absolute Min. -28.35 -20.42 -13.70 -6.05 2.25 5.52 9.64 9.64 1.00 -4.16 14.95 -23.32 -28.35 57% Wet Bulb Value 28.17 CPS/USAR CHAPTER 02 2.3-43 REV. 11, JANUARY 2005 TABLE 2.3-9 33 FOOT DEW POINT AVERAGE AVERAGE AVERAGE DAILY DAILY ABSOLUTE ABSOLUTE DAILY MAXIMUM MINIMUM MAXIMUM MINIMUM January -7.8 -4.4 -11.1 14.1 -29.5 February -4.0 -0.7 -7.5 13.6 -24.1 March 1.8 5.4 -1.2 17.7 -17.8 April 4.2 7.4 1.3 19.0 -10.0 May 8.1 11.0 5.2 22.7 -9.0 June 13.5 16.4 10.6 25.6 -0.3 July 16.5 19.3 14.0 25. 3.5 August 15.9 18.1 13.6 24.5 2.5 September 11.4 14.0 8.5 23.3 -7.1 October 4.2 7.1 1.4 9.1 -11.3 November -0.1 2.8 -2.7 16.3 -17.5 December -5.2 -2.1 -8.3 13.1 -25.7 Period of Record 4.7 7.8 1.9 25.6 -29.5 CPS/USAR CHAPTER 02 2.3-44 REV. 11, JANUARY 2005 TABLE 2.3-10 PERCENT OF HOURS WITH DEW POINT 18.3 DEG. C 12.8 DEG. C 7.2 DEG. C 0.0 DEG. C OR MORE OR MORE OR MORE OR MORE January 0.0 0.1 2.0 16.5 February 0.0 0.2 3.5 27.9 March 0.0 5.9 21.7 58.9 April 0.1 9.9 32.8 73.7 May 3.0 22.1 59.1 89.5 June 19.3 54.1 89.0 99.9 July 38.1 79.3 98.1 100.0 August 37.7 73.9 94.3 100.0 September 20.3 41.1 73.0 96.2 October 0.4 13.5 34.1 72.5 November 0.0 4.6 15.0 47.3 December 0.0 0.1 2.5 17.9 Period of Record 9.5 24.9 43.3 66.3 CPS/USAR CHAPTER 02 2.3-45 REV. 11, JANUARY 2005 TABLE 2.3-11 PERCENT OF HOURS WITH DEW POINT SPREAD 0.0 to 0.7 0.8 TO 2.2 2.3 TO 4.4 4.5 or DEG. C DEG. C DEG. C DEG. C January 15.8 33.0 37.3 14.0 February 20.1 20.7 26.8 32.3 March 6.6 18.0 29.0 46.5 April 3.4 14.2 21.1 61.2 May 1.4 9.0 22.7 66.9 June 3.0 11.1 20.5 65.4 July 2.6 8.3 22.0 67.1 August 3.0 16.3 25.9 54.8 September 5.0 16.8 23.5 54.7 October 4.5 14.9 16.2 64.4 November 7.6 20.8 31.1 40.6 December 12.7 26.7 31.8 18.8 Period of Record 7.0 18.4 25.8 48.8

CPS/USAR CHAPTER 02 2.3-46 REV. 11, JANUARY 2005 TABLE 2.3-12 PRECIPITATION (INCHES)

% HRS. WITH PRECIPITATION % DAYS WITH PRECIPITATION MAX. CONSECUTIVE HRS. MAX. CONSECUTIVE DAYS TOTAL MAXIMUM FOR 1 HR.

MAXIMUM IN. 1 DAY 0.01 IN.

OR MORE 1.00 IN. OR MORE 0.01 OR MORE 1.00 OR MORE WITH PRECIP. WITHOUT PRECIP. WITH PRECIP. WITHOUT PRECIP. January 7.00 0.50 2.53 3.4 0.0 21.3 0.6 14 356 5 14 February 5.74 0.26 0.97 3.3 0.0 19.9 0.0 9 470 3 19 March 17.21 0.69 1.29 5.9 0.0 23.3 1.9 10 408 3 16 April 9.16 0.69 1.63 3.4 0.0 25.1 0.6 14 455 5 18 May 8.98 0.52 0.62 3.6 0.0 26.0 0.0 6 293 5 12 June 20.80 1.15 2.72 4.7 0.0 31.3 3.3 14 545 5 22 July 11.34 0.43 1.74 3.1 0.0 25.2 0.6 7 365 4 14 August 12.59 0.80 1.34 2.9 0.0 21.9 0.6 8 476 3 21 September 12.20 0.81 1.26 3.8 0.0 28.0 2.0 11 372 8 15 October 7.64 0.45 0.94 3.7 0.0 20.6 0.0 12 332 3 13 November 9.13 0.40 1.06 4.4 0.0 22.0 0.7 11 620 5 25 December 6.67 0.34 0.93 3.7 0.0 21.9 0.0 8 406 8 16 Period of Record 128.45 1.15 2.72 3.8 0.0 24.6 0.9 14 807 8 33 CPS/USAR CHAPTER 02 2.3-47 REV. 11, JANUARY 2005 TABLE 2.3-13 MONTHLY FREQUENCY OF FOG OCCURRENCE, HOURS OF MAXIMUM AND MINIMUM, AND FOG PERSISTENCE FOR PEORIA, ILLINOIS (1949-1951; 1957-1971)

NUMBER OF TIMES IN 15 YEARS TOTAL FREQUENCY DAILY MAXIMUM DAILY MINIMUM FOG PERSISTED FOR AT LEAST MONTH OF OCCURRENCES (%) HOUR  % HOUR  % 12 HOURS 24 HOURS MAX. Jan. 17.8 8AM 25.1 6PM 14.0 38 15 95 Feb. 17.1 8AM 26.8 3PM 11.6 32 8 42 March 14.9 6AM 24.1 3PM 9.5 33 8 74 April 8.2 6AM 18.0 2PM 4.1 10 4 36 May 7.4 6AM 17.2 5PM 2.5 11 2 34 June 5.7 5AM 17.4 6PM 0.9 3 1 42 July 7.3 5AM 27.6 5PM 0.7 7 0 15 Aug. 8.6 6AM 35.7 4PM 0.4 5 0 19 Sept. 9.1 6AM 27.3 2PM 1.9 10 1 33 Oct. 10.3 7AM 23.3 3PM 5.4 15 3 34 Nov. 13.8 8AM 23.0 1PM 8.5 25 7 43 Dec. 15.5 9AM 21.5 4PM 10.0 38 9 48 CPS/USAR CHAPTER 02 2.3-48 REV. 11, JANUARY 2005 TABLE 2.3-14 MONTHLY FREQUENCY OF FOG OCCURRENCE, HOURS OF MAXIMUM AND MINIMUM, AND FOG PERSISTENCE FOR SPRINGFIELD, ILLINOIS (1951-1961; 1963-1970)

NUMBER OF TIMES IN 15 YEARS TOTAL FREQUENCY DAILY MAXIMUM DAILY MINIMUM FOG PERSISTED FOR AT LEAST MONTH OF OCCURRENCES (%) HOUR  % HOUR % 12 HOURS 24 HOURS MAX.January 17.2 7AM 25.1 3PM 13.4 49 17 90 February 15.0 7AM 23.9 3PM 10.8 39 15 53 March 12.7 6AM 21.4 3PM 8.7 36 8 36 April 6.4 6AM 16.1 4PM 2.3 16 2 26 May 5.5 5AM 14.6 4PM 1.5 8 1 27 June 3.7 6AM 12.4 5PM 0.8 1 1 29 July 5.0 5AM 22.3 3PM 0.2 6 0 19 August 6.1 6AM 27.0 4PM 0.2 2 0 13 September 5.5 6AM 23.9 4PM 0.3 3 0 22 October 6.7 6AM 15.8 4PM 4.0 14 3 47 November 9.4 7AM 17.4 2PM 4.9 25 5 51 December 15.4 8AM 20.8 2PM 12.2 37 17 75

CPS/USAR CHAPTER 02 2.3-49 REV. 11, JANUARY 2005 TABLE 2.3-15 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY A - DELTA T LESS THAN -1.8 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3- 1.4 1 4 3 2 2 7 9 5 5 6 2 3 4 3 4 5 65 (1) 0.06 0.23 0.17 0.11 0.11 0.40 0.51 0.28 0.28 0.34 0.11 0.17 0.23 0.17 0.23 0.28 3.68 (2) 0.00 0.01 0.01 0.00 0.00 0.02 0.02 0.01 0.01 0.01 0.00 0.01 0.01 0.01 0.01 0.05 0.16 1.5- 3.0 23 24 12 14 8 19 34 41 31 37 13 24 30 27 18 24 379 (1) 1.30 1.36 0.68 0.79 0.45 1.08 1.93 2.32 1.76 2.10 0.74 1.36 1.70 1.53 1.02 1.36 21.46 (2) 0.06 0.06 0.03 0.03 0.02 0.05 0.08 0.10 0.08 0.09 0.03 0.06 0.07 0.07 0.04 0.06 0.93 3.1- 5.0 39 43 26 19 8 17 38 61 40 65 32 44 37 57 24 29 579 (1) 2.21 2.43 1.47 1.08 0.45 0.96 2.15 3.45 2.27 3.68 1.81 2.49 2.10 3.23 1.36 1.64 32.79 (2) 0.10 0.11 0.06 0.05 0.02 0.04 0.09 0.15 0.10 0.16 0.08 0.11 0.09 0.14 0.06 0.07 1.42 5.1- 8.0 28 59 27 8 4 10 22 46 38 52 46 71 65 48 49 26 594 (1) 1.59 3.34 1.25 0.45 0.23 0.57 1.25 2.60 2.15 2.94 2.60 4.02 3.68 2.72 2.77 1.47 33.64 (2) 0.07 0.15 0.05 0.02 0.01 0.02 0.05 0.11 0.09 0.13 0.11 0.17 0.16 0.12 0.12 0.06 1.46 8.1-10.4 4 2 2 0 0 0 1 9 6 11 13 19 8 5 13 6 104 (1) 0.23 0.11 0.11 0.00 0.00 0.00 0.06 0.51 0.34 0.62 1.02 1.08 0.45 0.28 0.74 0.34 5.89 (2) 0.01 0.00 0.00 0.00 0.00 0.00 0.00 0.02 0.01 0.03 0.04 0.05 0.02 0.01 0.03 0.01 0.26 OVER 10.4 0 12 1 1 2 0 1 0 2 2 3 7 2 4 2 5 44 (1) 0.00 0.68 0.06 0.06 0.11 0.00 0.06 0.00 0.11 0.11 0.17 0.40 0.11 0.23 0.11 0.28 2.49 (2) 0.00 0.03 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.01 0.02 0.00 0.01 0.00 0.01 0.11 ALL SPEEDS 95 144 66 44 24 53 105 162 122 173 114 168 146 144 110 95 1765 (1) 5.38 8.15 3.74 2.49 1.36 3.00 5.95 9.17 6.91 9.80 6.46 9.51 8.27 8.15 6.23 5.38 99.94 (2) 0.23 0.35 0.16 0.11 0.06 0.13 0.26 0.40 0.30 0.43 0.28 0.41 0.36 0.35 0.27 0.23 4.34

(1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 1766 HRS ON THIS PAGE 1 HRS ( 0.1 PCT ) LESS THAN 0.3 MPS ( 0.0 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-50 REV. 11, JANUARY 2005 TABLE 2.3-16 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY B - DELTA T -1.8 TO -1.7 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3- 1.4 0 4 5 1 0 1 1 2 1 6 2 5 4 2 2 0 36 (1) 0.00 0.27 0.34 0.07 0.00 0.07 0.07 0.14 0.07 0.41 0.14 0.34 0.27 0.14 0.14 0.00 2.47 (2) 0.00 0.01 0.01 0.00 0.00 0.00 0.00 0.00 0.00 0.01 0.00 0.01 0.01 0.00 0.00 0.00 0.09 1.5- 3.0 12 24 8 13 10 10 14 22 13 36 22 15 18 15 13 15 260 (1) 0.82 1.65 0.55 0.89 0.69 0.69 0.96 1.51 0.69 2.47 1.51 1.03 1.24 1.03 0.89 1.03 17.86 (2) 0.03 0.06 0.02 0.03 0.02 0.02 0.03 0.05 0.03 0.09 0.05 0.04 0.04 0.04 0.03 0.04 0.64 3.1- 5.0 35 32 18 14 17 24 29 41 45 61 40 46 40 43 28 27 541 (1) 2.40 2.20 1.24 0.96 1.17 1.72 1.99 2.82 3.09 4.19 2.75 3.16 2.75 2.95 1.92 1.85 37.16 (2) 0.09 0.08 0.04 0.03 0.04 0.06 0.07 0.10 0.11 0.15 0.10 0.11 0.10 0.11 0.07 0.07 1.33 5.1- 8.0 20 34 16 20 6 16 31 27 35 46 42 40 47 47 22 26 475 (1) 1.37 2.34 1.10 1.37 0.41 1.10 2.13 1.85 2.40 3.16 2.88 2.76 3.23 3.23 1.51 1.79 32.62 (2) 0.05 0.08 0.04 0.05 0.01 0.04 0.08 0.07 0.09 0.11 0.10 0.10 0.12 0.12 0.05 0.06 1.17 8.1-10.4 3 0 0 1 0 0 2 7 5 5 9 24 16 4 3 3 82 (1) 0.21 0.00 0.00 0.07 0.00 0.00 0.14 0.48 0.34 0.34 0.62 1.65 1.10 0.27 0.21 0.21 5.63 (2) 0.01 0.00 0.00 0.00 0.00 0.00 0.00 0.02 0.01 0.01 0.02 0.06 0.04 0.01 0.01 0.01 0.20 OVER 10.4 2 1 0 2 6 2 1 6 3 4 5 8 15 1 0 5 61 (1) 0.14 0.07 0.00 0.14 0.41 0.14 0.07 0.41 0.21 0.27 0.34 0.55 1.03 0.07 0.00 0.34 4.19 (2) 0.00 0.00 0.00 0.00 0.01 0.00 0.00 0.01 0.01 0.01 0.01 0.02 0.04 0.00 0.00 0.01 0.15 ALL SPEEDS 72 95 47 51 39 54 78 105 102 158 120 138 140 112 68 76 1455 (1) 4.95 6.52 3.23 3.50 2.68 3.71 5.36 7.21 7.01 10.85 8.24 9.48 9.62 7.69 4.67 5.22 99.93 (2) 0.18 0.23 0.12 0.13 0.10 0.13 0.19 0.26 0.25 0.39 0.30 0.34 0.34 0.28 0.17 0.19 3.58 (1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 1456 HRS ON THIS PAGE 1 HRS ( 0.1 PCT ) LESS THAN 0.3 MPS ( 0.0 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-51 REV. 11, JANUARY 2005 TABLE 2.3-17 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY C - DELTA T -1.6 TO -1.5 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3 - 1.4 0 5 4 1 1 3 7 7 7 4 5 5 6 4 3 2 64 (1) 0.00 0.23 0.18 0.05 0.05 0.14 0.32 0.32 0.32 0.18 0.23 0.23 0.27 0.18 0.14 0.09 2.92 (2) 0.00 0.01 0.01 0.00 0.00 0.01 0.02 0.02 0.02 0.01 0.01 0.01 0.01 0.01 0.01 0.00 0.16 1.5- 3.0 27 31 31 18 12 25 29 36 29 32 22 28 35 18 28 22 423 (1) 1.23 1.42 1.42 0.82 0.55 1.14 1.32 1.64 1.32 1.46 1.01 1.28 1.60 0.82 1.28 1.01 19.32 (2) 0.07 0.08 0.08 0.04 0.03 0.06 0.07 0.09 0.07 0.08 0.05 0.07 0.09 0.04 0.07 0.05 1.04 3.1- 5.0 42 46 40 31 31 24 51 55 47 83 67 38 62 50 52 27 746 (1) 1.92 2.10 1.83 1.42 1.42 1.10 2.33 2.51 2.15 3.79 3.06 1.74 2.83 2.28 2.38 1.23 34.08 (2) 0.10 0.11 0.10 0.08 0.08 0.06 0.13 0.14 0.12 0.20 0.16 0.09 0.15 0.12 0.13 0.07 1.83 5.1- 8.0 35 34 19 20 20 31 40 33 43 88 62 61 72 55 33 29 675 (1) 1.60 1.55 0.87 0.91 0.91 1.42 1.83 1.51 1.96 4.02 2.83 2.79 3.29 2.51 1.51 1.32 30.84 (2) 0.09 0.08 0.05 0.05 0.05 0.08 0.10 0.08 0.11 0.22 0.15 0.15 0.18 0.14 0.08 0.07 1.66 8.1-10.4 8 3 0 1 0 2 2 9 14 12 17 36 20 13 5 7 149 (1) 0.37 0.14 0.00 0.05 0.00 0.09 0.09 0.41 0.64 0.55 0.78 1.64 0.91 0.59 0.23 0.32 6.81 (2) 0.02 0.01 0.00 0.00 0.00 0.00 0.00 0.02 0.03 0.03 0.04 0.09 0.05 0.03 0.01 0.02 0.37 OVER 10.4 1 3 1 8 7 9 10 3 12 9 19 23 12 4 4 5 130 (1) 0.05 0.14 0.05 0.37 0.32 0.41 0.46 0.14 0.55 0.41 0.87 1.05 0.55 0.18 0.18 0.23 5.94 (2) 0.00 0.01 0.00 0.02 0.02 0.02 0.02 0.01 0.03 0.02 0.05 0.06 0.03 0.01 0.01 0.01 0.32 ALL SPEEDS 113 122 95 79 71 94 139 143 152 228 192 191 207 144 125 92 2187 (1) 5.16 5.57 4.34 3.61 3.24 4.29 6.35 6.53 6.94 10.42 8.77 8.73 9.46 6.58 5.71 4.20 99.91 (2) 0.28 0.30 0.23 0.19 0.17 0.23 0.34 0.35 0.37 0.56 0.47 0.47 0.51 0.35 0.31 0.23 5.38

(1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 2189 HRS ON THIS PAGE 2 HRS ( 0.1 PCT ) LESS THAN 0.3 MPS ( 0.0 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-52 REV. 11, JANUARY 2005 TABLE 2.3-18 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY D - DELTA T -1.4 TO -0.5 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3 - 1.4 30 34 31 37 40 25 46 50 46 52 37 36 46 26 35 31 602 (1) 0.18 0.21 0.19 0.23 0.25 0.15 0.28 0.31 0.28 0.32 0.23 0.22 0.28 0.16 0.21 0.19 3.69 (2) 0.07 0.08 0.08 0.09 0.10 0.06 0.11 0.12 0.11 0.13 0.09 0.09 0.11 0.06 0.09 0.08 1.48 1.5- 3.0 126 178 204 197 147 173 250 249 218 229 160 162 190 166 155 135 2939 (1) 0.77 1.09 1.25 1.21 0.90 1.06 1.53 1.53 1.34 1.40 0.98 0.99 1.16 1.02 0.95 0.83 18.01 (2) 0.31 0.44 0.50 0.48 0.36 0.43 0.61 0.61 0.54 0.56 0.39 0.40 0.47 0.41 0.38 0.33 7.23 3.1- 5.0 269 289 291 286 248 231 302 416 466 396 314 360 450 406 316 294 5334 (1) 1.65 1.77 1.78 1.75 1.52 1.42 1.85 2.55 2.86 2.43 1.92 2.21 2.76 2.49 1.94 1.80 32.69 (2) 0.66 0.71 0.72 0.70 0.61 0.57 0.74 1.02 1.15 0.97 0.77 0.89 1.11 1.00 0.78 0.72 13.11 5.1- 8.0 240 263 138 134 170 193 228 439 515 428 323 535 679 457 319 269 5330 (1) 1.47 1.61 0.85 0.82 1.04 1.18 1.40 2.69 3.16 2.62 1.98 3.28 4.16 2.80 1.96 1.65 32.67 (2) 0.59 0.65 0.34 0.33 0.42 0.47 0.56 1.08 1.27 1.05 0.79 1.32 1.67 1.12 0.78 0.66 13.10 8.1-10.4 65 63 11 16 16 23 40 152 139 119 137 200 204 102 86 73 1446 (1) 0.40 0.39 0.07 0.10 0.10 0.14 0.25 0.93 0.85 0.73 0.84 1.23 1.25 0.63 0.53 0.85 8.86 (2) 0.16 0.15 0.03 0.04 0.04 0.06 0.10 0.37 0.34 0.29 0.34 0.42 0.50 0.25 0.21 0.18 3.55 OVER 10.4 25 19 13 21 18 22 17 39 58 52 95 132 80 24 24 23 662 (1) 0.15 0.12 0.08 0.13 0.11 0.13 0.10 0.24 0.36 0.32 0.58 0.81 0.49 0.15 0.15 0.14 4.06 (2) 0.06 0.05 0.03 0.05 0.04 0.05 0.04 0.10 0.14 0.13 0.23 0.32 0.20 0.06 0.06 0.06 1.63 ALL SPEEDS 755 846 688 691 639 667 883 1345 1442 1276 1066 1425 1649 1181 935 825 16313 (1) 4.63 5.18 4.22 4.23 3.92 4.09 5.41 8.24 8.84 7.82 6.53 8.73 10.11 7.24 5.73 5.06 99.98 (2) 1.86 2.08 1.69 1.70 1.57 1.64 26.17 3.31 3.55 3.14 2.62 3.50 4.05 2.90 2.30 2.03 40.10 (1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 16317 HRS ON THIS PAGE 4 HRS ( 0.0 PCT ) LESS THAN 0.3 MPS ( 0.0 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-53 REV. 11, JANUARY 2005 TABLE 2.3-19 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY E - DELTA T -0.4 TO +1.5 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3 - 1.4 38 42 49 47 33 53 62 69 60 60 48 44 41 28 19 32 725 (1) 0.35 0.39 0.45 0.43 0.31 0.49 0.57 0.64 0.55 0.55 0.44 0.41 0.38 0.26 0.15 0.30 6.70 (2) 0.09 0.10 0.12 0.12 0.08 0.13 0.15 0.17 0.15 0.15 0.12 0.11 0.10 0.07 0.05 0.08 1.78 1.5- 3.0 95 170 188 204 201 255 308 312 299 218 197 173 175 159 113 98 3165 (1) 0.88 1.57 1.74 1.89 1.86 2.36 2.85 2.88 2.76 2.02 1.82 1.60 1.62 1.47 1.04 0.91 29.26 (2) 0.23 0.42 0.46 0.50 0.49 0.63 0.76 0.77 0.74 0.54 0.48 0.43 0.43 0.39 0.28 0.24 7.78 3.1- 5.0 119 156 162 187 197 246 367 530 518 343 241 242 223 148 116 151 3946 (1) 1.10 1.44 1.50 1.73 1.82 2.27 3.39 4.90 4.79 3.17 2.23 2.24 2.06 1.37 1.07 1.40 36.49 (2) 0.29 0.38 0.40 0.46 0.48 0.60 0.90 1.30 1.27 0.84 0.59 0.59 0.55 0.36 0.29 0.37 9.70 5.1- 8.0 48 72 33 56 100 148 174 402 386 193 188 197 124 56 42 65 2284 (1) 0.44 0.67 0.31 0.52 0.92 1.37 1.61 3.72 3.57 1.78 1.74 1.82 1.15 0.52 0.39 0.60 21.12 (2) 0.12 0.18 0.08 0.14 0.25 0.36 0.43 0.99 0.95 0.47 0.46 0.48 0.30 0.14 0.10 0.16 5.61 8.1-10.4 15 10 5 2 21 26 19 56 43 32 46 51 25 9 20 14 394 (1) 0.14 0.09 0.05 0.02 0.19 0.24 0.18 0.52 0.40 0.30 0.43 0.47 0.23 0.08 0.18 0.13 3.64 (2) 0.04 0.02 0.01 0.00 0.05 0.06 0.05 0.14 0.11 0.08 0.11 0.13 0.06 0.02 0.05 0.03 0.97 OVER 10.4 4 9 9 17 24 15 20 31 36 24 24 23 13 13 4 9 275 (1) 0.04 0.08 0.08 0.16 0.22 0.14 0.18 0.29 0.33 0.22 0.22 0.21 0.12 0.12 0.04 0.08 2.54 (2) 0.01 0.02 0.02 0.04 0.06 0.04 0.05 0.08 0.09 0.06 0.06 0.06 0.03 0.03 0.01 0.02 0.68 ALL SPEEDS 319 459 446 513 576 743 950 1480 1342 870 744 730 601 413 314 369 10789 (1) 2.95 4.24 4.12 4.74 5.33 6.87 8.78 12.94 12.41 8.04 6.88 6.75 5.56 3.82 2.90 3.41 99.76 (2) 0.78 1.13 1.10 1.26 1.42 1.83 2.34 3.44 3.30 2.14 1.83 1.79 1.48 1.02 0.77 0.91 26.52 (1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 10815 HRS ON THIS PAGE 26 HRS ( 0.2 PCT ) LESS THAN 0.3 MPS ( 0.1 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-54 REV. 11, JANUARY 2005 TABLE 2.3-20 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY F - DELTA T 1.6 TO 4.0 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3 - 1.4 30 50 50 42 36 49 54 59 36 44 35 44 29 25 33 39 655 (1) 0.67 1.12 1.12 0.94 0.80 1.10 1.21 1.32 0.80 0.98 0.78 0.98 0.65 0.56 0.74 0.87 14.64 (2) 0.07 0.12 0.12 0.10 0.09 0.12 0.13 0.15 0.09 0.11 0.09 0.11 0.07 0.06 0.08 0.10 1.61 1.5- 3.0 75 125 134 153 161 197 216 222 248 209 152 139 163 113 63 83 2453 (1) 1.68 2.79 3.00 3.42 3.60 4.40 4.83 4.96 5.54 4.67 3.40 3.11 3.64 2.53 1.41 1.86 54.83 (2) 0.18 0.31 0.33 0.38 0.40 0.48 0.53 0.55 0.61 0.51 0.37 0.34 0.40 0.28 0.15 0.20 6.03 3.1- 5.0 26 24 22 28 40 56 101 114 148 120 96 73 75 57 24 27 1031 (1) 0.58 0.54 0.49 0.63 0.89 1.25 2.26 2.55 3.31 2.68 2.15 1.63 1.68 1.27 0.54 0.60 23.04 (2) 0.06 0.06 0.05 0.07 0.10 0.14 0.25 0.28 0.36 0.30 0.24 0.18 0.18 0.14 0.06 0.07 2.53 5.1- 8.0 0 0 0 0 0 5 4 4 8 14 10 16 10 3 4 2 80 (1) 0.00 0.00 0.00 0.00 0.00 0.11 0.09 0.09 0.18 0.31 0.22 0.36 0.22 0.07 0.09 0.04 1.79 (2) 0.00 0.00 0.00 0.00 0.00 0.01 0.01 0.01 0.02 0.03 0.02 0.04 0.02 0.01 0.01 0.00 0.20 8.1-10.4 0 0 0 0 0 0 0 0 0 0 0 1 1 1 2 0 5 (1) 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.02 0.02 0.02 0.04 0.00 0.11 (2) 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.01 OVER 10.4 11 21 14 22 9 13 23 18 23 17 15 12 8 5 4 9 224 (1) 0.25 0.47 0.31 0.49 0.20 0.29 0.51 0.40 0.51 0.38 0.34 0.27 0.18 0.11 0.09 0.20 5.01 (2) 0.03 0.05 0.03 0.05 0.02 0.03 0.06 0.04 0.06 0.04 0.04 0.03 0.02 0.01 0.01 0.02 0.55 ALL SPEEDS 142 220 220 245 246 320 398 417 463 404 308 285 286 204 130 160 4448 (1) 3.17 4.92 4.92 5.48 5.50 7.15 8.90 9.32 10.35 9.03 6.88 6.37 6.39 4.56 2.91 3.58 99.42 (2) 0.35 0.54 0.54 0.60 0.60 0.79 0.98 1.03 1.14 0.99 0.76 0.70 0.70 0.50 0.32 0.39 10.93 (1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 4474 HRS ON THIS PAGE 24 HRS ( 0.6 PCT ) LESS THAN 0.3 MPS ( 0.1 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-55 REV. 11, JANUARY 2005 TABLE 2.3-21 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T STABILITY G - DELTA T GREATER THAN 4.0 DEG C PER 100 METERS DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3 - 1.4 53 73 73 79 52 57 69 98 78 63 58 58 55 49 41 37 993 (1) 1.45 1.99 1.99 2.16 1.42 1.56 1.89 2.68 2.13 1.72 1.58 1.58 1.50 1.34 1.12 1.01 27.13 (2) 0.13 0.18 0.18 0.19 0.13 0.14 0.17 0.24 0.19 0.15 0.14 0.14 0.14 0.12 0.10 0.09 2.44 1.5- 3.0 75 138 94 93 90 160 182 189 216 151 88 94 92 96 43 57 1858 (1) 2.05 3.77 2.57 2.54 2.46 4.37 4.97 5.16 5.90 4.13 2.40 2.57 2.51 2.62 1.17 1.56 50.77 (2) 0.18 0.34 0.23 0.23 0.22 0.39 0.45 0.46 0.53 0.37 0.22 0.23 0.23 0.24 0.11 0.14 4.57 3.1- 5.0 8 9 9 10 13 19 23 23 55 28 13 17 22 27 12 7 295 (1) 0.22 0.25 0.25 0.27 0.36 0.52 0.63 0.63 1.50 0.77 0.36 0.46 0.60 0.74 0.33 0.19 8.06 (2) 0.02 0.02 0.02 0.02 0.03 0.05 0.06 0.06 0.14 0.07 0.03 0.04 0.05 0.07 0.03 0.02 0.73 5.1- 8.0 6 10 1 5 14 15 4 35 55 13 2 17 14 2 1 3 197 (1) 0.16 0.27 0.03 0.14 0.38 0.41 0.11 0.96 1.50 0.36 0.05 0.46 0.38 0.05 0.03 0.08 5.38 (2) 0.01 0.02 0.00 0.01 0.03 0.04 0.01 0.09 0.14 0.03 0.00 0.04 0.03 0.00 0.00 0.01 0.48 8.1-10.4 1 1 1 0 0 0 0 0 20 4 1 8 6 0 2 3 47 (1) 0.03 0.03 0.03 0.00 0.00 0.00 0.00 0.00 0.55 0.11 0.03 0.22 0.16 0.00 0.05 0.08 1.28 (2) 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.05 0.01 0.00 0.02 0.01 0.00 0.00 0.01 0.12 OVER 10.4 8 30 27 25 15 9 16 27 16 13 16 2 5 5 2 5 221 (1) 0.22 0.82 0.74 0.68 0.41 0.25 0.44 0.74 0.44 0.36 0.44 0.05 0.14 0.14 0.05 0.14 6.04 (2) 0.02 0.07 0.07 0.06 0.04 0.02 0.04 0.07 0.04 0.03 0.04 0.00 0.01 0.01 0.00 0.01 0.54 ALL SPEEDS 151 261 205 212 184 260 294 372 440 272 178 196 194 179 101 112 3611 (1) 4.13 7.13 5.60 5.79 5.03 7.10 8.03 10.16 12.02 7.43 4.86 5.36 5.30 4.89 2.76 3.06 98.66 (2) 0.37 0.64 0.50 0.52 0.45 0.64 0.72 0.91 1.08 0.67 0.44 0.48 0.48 0.44 0.25 0.28 8.88 (1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 3660 HRS ON THIS PAGE 49 HRS ( 1.3 PCT ) LESS THAN 0.3 MPS ( 0.1 PCT OF ALL HRS)

CPS/USAR CHAPTER 02 2.3-56 REV. 11, JANUARY 2005 TABLE 2.3-22 JOINT FREQUENCY DISTRIBUTION CLINTON POWER STATION 33 FT WIND DISTRIBUTION OF WIND DIRECTIONS AND SPEEDS 4/14/72 - 4/30/77 198-33 FT DELTA T ALL STABILITIES COMBINED DIRECTION SPEED (MPS) NNE NE ENE E ESE SE SSE S SSW SW WSW W WNW NW NNW N TOTAL 0.3 - 1.4 152 212 215 209 164 195 248 290 233 235 187 195 185 137 137 146 3140 (1) 0.37 0.52 0.53 0.51 0.40 0.48 0.61 0.71 0.57 0.58 0.46 0.48 0.45 0.34 0.34 0.36 7.72 (2) 0.37 0.52 0.53 0.51 0.40 0.48 0.61 0.71 0.57 0.58 0.46 0.48 0.45 0.34 0.34 0.36 7.72 1.5- 3.0 433 690 671 692 629 839 1033 1071 1054 912 654 635 703 594 433 434 11477 (1) 1.06 1.70 1.65 1.70 1.55 2.06 2.54 2.63 2.59 2.24 1.61 1.56 1.73 1.46 1.06 1.07 28.21 (2) 1.06 1.70 1.65 1.70 1.55 2.06 2.54 2.63 2.59 2.24 1.61 1.56 1.73 1.46 1.06 1.07 28.21 3.1- 5.0 538 599 568 575 554 618 911 1240 1319 1096 803 820 909 788 572 562 12472 (1) 1.32 1.47 1.40 1.41 1.36 1.52 2.24 3.05 3.24 2.69 1.97 2.02 2.23 1.94 1.41 1.38 30.66 (2) 1.32 1.47 1.40 1.41 1.36 1.52 2.24 3.05 3.24 2.69 1.97 2.02 2.23 1.94 1.41 1.38 30.66 5.1- 8.0 377 472 229 243 314 418 503 956 1000 834 673 937 1011 668 470 420 9635 (1) 0.93 1.16 0.56 0.60 0.77 1.03 1.24 2.42 2.66 2.05 1.65 2.30 2.49 1.64 1.16 1.03 23.69 (2) 0.93 1.16 0.56 0.60 0.77 1.03 1.24 2.42 2.66 2.05 1.65 2.30 2.49 1.64 1.16 1.03 23.69 8.1-10.4 96 79 19 20 37 51 64 233 227 183 228 339 280 134 131 106 2227 (1) 0.24 0.19 0.05 0.05 0.09 0.13 0.16 0.57 0.56 0.45 0.56 0.83 0.69 0.33 0.32 0.26 5.47 (2) 0.24 0.19 0.05 0.05 0.09 0.13 0.16 0.57 0.56 0.45 0.56 0.83 0.69 0.33 0.32 0.26 5.47 OVER 10.4 51 95 65 96 81 70 88 124 150 121 177 207 135 56 40 61 1617 (1) 0.13 0.23 0.16 0.24 0.20 0.17 0.22 0.30 0.37 0.30 0.44 0.51 0.33 0.14 0.10 0.15 3.98 (2) 0.13 0.23 0.16 0.24 0.20 0.17 0.22 0.30 0.37 0.30 0.44 0.51 0.33 0.14 0.10 0.15 3.98 ALL SPEEDS 1647 2147 1767 18 35 1779 2191 2847 3944 4063 3381 2722 3133 3223 2377 1783 1729 40568 (1) 4.05 5.28 4.34 4.51 4.37 5.39 7.00 9.70 9.99 8.31 6.69 7.70 7.92 5.84 4.38 4.25 99.73 (2) 4.05 5.28 4.34 4.51 4.37 5.39 7.00 9.70 9.99 8.31 6.69 7.70 7.92 5.84 4.38 4.25 99.73

(1) = PERCENT OF ALL GOOD OB S FOR THIS PAGE (2) = PERCENT OF ALL GOOD OB S FOR THE PERIOD 40677 GOOD HRS 109 HRS ( 0.3 PCT ) LESS THAN 0.3 MPS 44208 HRS IN THE TIME PERIOD 92.0 PCT DATA RECOVERY

CPS/USAR CHAPTER 02 2.3-57 REV. 11, JANUARY 2005 TABLE 2.3-23 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 1)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE MONTH OR LESS OR LESS OR LESS OR LESS January 218 299 395 419 February 120 235 349 360 March 108 151 181 195 April 25 77 151 109 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 60 93 134 144 September 77 131 157 173 October 84 133 231 288 November 133 342 399 430 December 227 416 453 465

• Fog prediction model indicates minor steam fog for these months. Values shown are from Peoria, Illinois (1960-1970).

1/16 mile = 330 feet 3/16 mile = 990 feet CPS/USAR CHAPTER 02 2.3-58 REV. 11, JANUARY 2005 TABLE 2.3-24 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 2)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE OR LESS OR LESS OR LESS OR LESS January 108 163 163 163 February 44 148 152 152 March 41 72 76 76 April 0 33 44 46 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 93 99 103 111 September 80 97 107 107 October 72 102 105 109 November 44 135 149 149 December 60 139 139 148

• Fog prediction model indicates minor steam fog for these months. Values shown are from Peoria, Illinois (1960-1970).

1/16 mile = 330 feet

3/16 mile - 990 feet CPS/USAR CHAPTER 02 2.3-59 REV. 11, JANUARY 2005 TABLE 2.3-25 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 3)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE OR LESS OR LESS OR LESS OR LESS January 7 65 114 144 February 3 71 99 116 March 3 39 49 68 April 0 16 20 22 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 17 47 53 59 September 22 61 61 64 October 3 90 90 96 November 7 65 100 111 December 14 147 138 148

• Fog prediction model indicates minor steam fog for these months. Values shown are from Peoria, Illinois (1960-1970).

1/16 mile = 330 feet 3/16 mile = 990 feet CPS/USAR CHAPTER 02 2.3-60 REV. 11, JANUARY 2005 TABLE 2.3-26 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 4)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE OR LESS OR LESS OR LESS OR LESS January 7 7 17 50 February 3 3 5 34 March 3 3 4 28 April 0 0 1 8 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 0 0 1 6 September 1 7 8 14 October 3 10 14 21 November 4 4 9 26 December 2 2 6 24

• Fog prediction model indicates minor steam fog for these months. Values shown are from Peoria, Illinois (1960-1979).

1/16 mile = 330 feet

3/16 mile = 990 feet CPS/USAR CHAPTER 02 2.3-61 REV. 11, JANUARY 2005 TABLE 2.3-27 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 5)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE MONTH OR LESS OR LESS OR LESS OR LESS January 7 20 38 87 February 3 10 39 75 March 3 5 20 38 April 0 0 1 6 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 0 17 24 47 September 1 29 48 63 October 3 14 46 59 November 4 26 34 48 December 2 7 21 60

• Fog prediction model indicates minor steam fcg for these months. Values shown are from Peoria, Illinois (1960-1970).

1/16 mile = 330 feet

3/16 mile = 990 feet CPS/USAR CHAPTER 02 2.3-62 REV. 11, JANUARY 2005 TABLE 2.3-28 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 6)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE MONTH OR LESS OR LESS OR LESS OR LESS January 7 87 123 156 February 3 21 47 88 March 3 8 17 38 April 0 6 11 18 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 3 12 24 37 September 9 57 82 88 October 6 81 141 147 November 15 74 125 168 December 6 60 102 144

• Fog prediction model indicates minor steam fog for these months. Values shown are from Peoria, Illinois (1960-1970).

1/16 mile = 330 feet

3/16 mile = 990 feet CPS/USAR CHAPTER 02 2.3-63 REV. 11, JANUARY 2005 TABLE 2.3-29 PREDICTED NUMBER OF HOURS OF LAKE STEAM FOG (AREA 7)

VISIBILITY 100 FEET 1/16 MILE 3/16 MILE 1 MILE MONTH OR LESS OR LESS OR LESS OR LESS January 3 16 61 128 February 1 15 58 98 March 0 3 4 40 April 0 0 22 45 May* 1 1 1 5 June* 0 0 1 4 July* 1 1 2 7 August 0 20 22 50 September 4 23 33 52 October 4 22 28 39 November 2 7 23 79 December 1 16 61 138

• Fog prediction model indicates minor steam fog for these months. Values shown are from Peoria, Illinois (1960-1970).

1/16 mile = 330 feet

3/16 mile = 990 feet CPS/USAR CHAPTER 02 2.3-64 REV. 11, JANUARY 2005 TABLE 2.3-30 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 1 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTION (METERS) /Q /Q SSW 975 0.182E-03 0.274E-04 SW 975 0.190E-03 0.294E-04 WSW 975 0.210E-03 0.349E-04 W 975 0.211E-03 0.376E-04 WNW 975 0.169E-03 0.361E-04 NW 975 0.177E-03 0.377E-04 NNW 975 0.168E-03 0.350E-04 N 975 0.163E-03 0.291E-04 NNE 975 0.151E-03 0.311E-04 NE 975 0.154E-03 0.289E-04 ENE 975 0.153E-03 0.279E-04 E 975 0.150E-03 0.254E-04 ESE 975 0.143E-03 0.248E-04 SE 975 0.149E-03 0.258E-04 SSE 975 0.164E-03 0.254E-04 S 975 0.156E-03 0.277E-04 All Direction Case 0.178E-03 0.305E-04 Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-65 REV. 11, JANUARY 2005 TABLE 2.3-31 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 1 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4018 0.427E-04 0.347E-05 SW 4018 0.449E-04 0.379E-05 WSW 4018 0.475E-04 0.488E-05 W 4018 0.476E-04 0.528E-05 WNW 4018 0.379E-04 0.505E-05 NW 4018 0.401E-04 0.527E-05 NNW 4018 0.379E-04 0.473E-05 N 4018 0.342E-04 0.377E-05 NNE 4018 0.336E-04 0.425E-05 NE 4018 0.344E-04 0.374E-05 ENE 4018 0.354E-04 0.363E-05 E 4018 0.310E-04 0.315E-05 ESE 4018 0.282E-04 0.303E-05 SE 4018 0.331E-04 0.313E-05 SSE 4018 0.372E-04 0.304E-05 S 4018 0.367E-04 0.353E-05 All Direction Case 0.415E-04 0.426E-05

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-66 REV. 11, JANUARY 2005 TABLE 2.3-32 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 1 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4727 0.352E-04 0.271E-05 SW 5121 0.343E-04 0.270E-05 WSW 3482 0.568E-04 0.592E-05 W 2377 0.878E-04 0.110E-04 WNW 1508 0.115E-03 0.203E-04 NW 1585 0.114E-03 0.197E-04 NNW 1615 0.108E-03 0.174E-04 N 1585 0.105E-03 0.144E-04 NNE 1615 0.944E-04 0.156E-04 NE 1402 0.112E-03 0.174E-04 ENE 1189 0.127E-03 0.210E-04 E 1158 0.128E-03 0.197E-04 ESE 4724 0.232E-04 0.239E-05 SE 4077 0.328E-04 0.305E-05 SSE 3353 0.467E-04 0.395E-05 S 3353 0.453E-04 0.455E-05 Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-67 REV. 11, JANUARY 2005 TABLE 2.3-33 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 2 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 975 0.120E-03 0.193E-04 SW 975 0.137E-03 0.223E-04 WSW 975 0.141E-03 0.247E-04 W 975 0.141E-03 0.251E-04 WNW 975 0.118E-03 0.247E-04 NW 975 0.137E-03 0.247E-04 NNW 975 0.131E-03 0.241E-04 N 975 0.124E-03 0.214E-04 NNE 975 0.115E-03 0.226E-04 NE 975 0.113E-03 0.198E-04 ENE 975 0.101E-03 0.197E-04 E 975 0.982E-04 0.181E-04 ESE 975 0.945E-04 0.177E-04 SE 975 0.102E-03 0.173E-04 SSE 975 0.107E-03 0.169E-04 S 975 0.112E-03 0.200E-04 All Direction Case 0.126E-03 0.231E-04

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-68 REV. 11, JANUARY 2005 TABLE 2.3-34 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 2 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4018 0.284E-04 0.256E-05 SW 4018 0.315E-04 0.287E-05 WSW 4018 0.317E-04 0.346E-05 W 4018 0.305E-04 0.366E-05 WNW 4018 0.248E-04 0.356E-05 NW 4018 0.294E-04 0.357E-05 NNW 4018 0.266E-04 0.331E-05 N 4018 0.247E-04 0.279E-05 NNE 4018 0.246E-04 0.299E-05 NE 4018 0.247E-04 0.261E-05 ENE 4018 0.230E-04 0.264E-05 E 4018 0.217E-04 0.236E-05 ESE 4018 0.194E-04 0.229E-05 SE 4018 0.217E-04 0.220E-05 SSE 4018 0.234E-04 0.216E-05 S 4018 0.237E-04 0.264E-05 All Direction Case 0.272E-04 0.308E-05 Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-69 REV. 11, JANUARY 2005 TABLE 2.3-35 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 2 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4727 0.230E-04 0.203E-05 SW 5121 0.238E-04 0.206E-05 WSW 3482 0.388E-04 0.428E-05 W 2377 0.575E-04 0.768E-05 WNW 1508 0.774E-04 0.143E-04 NW 1585 0.862E-04 0.132E-04 NNW 1615 0.807E-04 0.124E-04 N 1585 0.763E-04 0.107E-04 NNE 1615 0.697E-04 0.112E-04 NE 1402 0.772E-04 0.119E-04 ENE 1189 0.813E-04 0.153E-04 E 1158 0.814E-04 0.142E-04 ESE 4724 0.157E-04 0.180E-05 SE 4077 0.214E-04 0.216E-05 SSE 3353 0.283E-04 0.283E-05 S 3353 0.296E-04 0.343E-05 Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-70 REV. 11, JANUARY 2005 TABLE 2.3-36 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 8 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 975 0.517E-04 0.891E-05 SW 975 0.660E-04 0.104E-04 WSW 975 0.606E-04 0.113E-04 W 975 0.647E-04 0.124E-04 WNW 975 0.529E-04 0.111E-04 NW 975 0.605E-04 0.111E-04 NNW 975 0.621E-04 0.111E-04 N 975 0.596E-04 0.108E-04 NNE 975 0.605E-04 0.102E-04 NE 975 0.548E-04 0.890E-05 ENE 975 0.489E-04 0.804E-05 E 975 0.464E-04 0.833E-05 ESE 975 0.490E-04 0.887E-05 SE 975 0.450E-04 0.836E-05 SSE 975 0.431E-04 0.734E-05 S 975 0.488E-04 0.890E-05 All Direction Case 0.600E-04 0.104E-04

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-71 REV. 11, JANUARY 2005 TABLE 2.3-37 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 8 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS /Q /Q SSW 4018 0.118E-04 0.123E-05 SW 4018 0.142E-04 0.147E-05 WSW 4018 0.129E-04 0.162E-05 W 4018 0.134E-04 0.179E-05 WNW 4018 0.104E-04 0.162E-05 NW 4018 0.125E-04 0.160E-05 NNW 4018 0.124E-04 0.155E-05 N 4018 0.118E-04 0.147E-05 NNE 4018 0.117E-04 0.139E-05 NE 4018 0.112E-04 0.121E-05 ENE 4018 0.964E-05 0.113E-05 E 4018 0.946E-05 0.115E-05 ESE 4018 0.100E-04 0.118E-05 SE 4018 0.931E-05 0.114E-05 SSE 4018 0.943E-05 0.101E-05 S 4018 0.921E-05 0.123E-05 All Direction Case 0.125E-04 0.147E-05

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-72 REV. 11, JANUARY 2005 TABLE 2.3-38 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 16 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 975 0.327E-04 0.588E-05 SW 975 0.403E-04 0.719E-05 WSW 975 0.396E-04 0.714E-05 W 975 0.434E-04 0.859E-05 WNW 975 0.332E-04 0.727E-05 NW 975 0.393E-04 0.725E-05 NNW 975 0.406E-04 0.753E-05 N 975 0.407E-04 0.771E-05 NNE 975 0.403E-04 0.693E-05 NE 975 0.380E-04 0.580E-05 ENE 975 0.320E-04 0.513E-05 E 975 0.312E-04 0.565E-05 ESE 975 0.342E-04 0.602E-05 SE 975 0.307E-04 0.537E-05 SSE 975 0.289E-04 0.469E-05 S 975 0.290E-04 0.584E-05 All Direction Case 0.403E-04 0.710E-05

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-73 REV. 11, JANUARY 2005 TABLE 2.3-39 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 16 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4018 0.712E-05 0.860E-06 SW 4018 0.869E-05 0.107E-05 WSW 4018 0.824E-05 0.105E-05 W 4018 0.905E-05 0.131E-05 WNW 4018 0.669E-05 0.112E-05 NW 4018 0.775E-05 0.109E-05 NNW 4018 0.764E-05 0.113E-05 N 4018 0.797E-05 0.111E-05 NNE 4018 0.770E-05 0.997E-06 NE 4018 0.758E-05 0.815E-06 ENE 4018 0.647E-05 0.736E-06 E 4018 0.661E-05 0.792E-06 ESE 4018 0.673E-05 0.841E-06 SE 4018 0.610E-05 0.740E-06 SSE 4018 0.596E-05 0.633E-06 S 4018 0.579E-05 0.810E-06 All Direction Case 0.820E-05 0.100E-05

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-74 REV. 11, JANUARY 2005 TABLE 2.3-40 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 72 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 975 0.125E-04 0.228E-05 SW 975 0.174E-04 0.318E-05 WSW 975 0.148E-04 0.303E-05 W 975 0.162E-04 0.350E-05 WNW 975 0.132E-04 0.305E-05 NW 975 0.151E-04 0.312E-05 NNW 975 0.181E-04 0.358E-05 N 975 0.185E-04 0.399E-05 NNE 975 0.182E-04 0.370E-05 NE 975 0.157E-04 0.307E-05 ENE 975 0.135E-04 0.244E-05 E 975 0.128E-04 0.269E-05 ESE 975 0.144E-04 0.269E-05 SE 975 0.136E-04 0.228E-05 SSE 975 0.123E-04 0.191E-05 S 975 0.130E-04 0.204E-05 All Direction Case 0.171E-04 0.320E-05

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-75 REV. 11, JANUARY 2005 TABLE 2.3-41 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 72 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4018 0.258E-05 0.360E-06 SW 4018 0.348E-05 0.478E-06 WSW 4018 0.317E-05 0.489E-06 W 4018 0.354E-05 0.551E-06 WNW 4018 0.248E-05 0.487E-06 NW 4018 0.292E-05 0.521E-06 NNW 4018 0.356E-05 0.541E-06 N 4018 0.343E-05 0.600E-06 NNE 4018 0.335E-05 0.575E-06 NE 4018 0.329E-05 0.457E-06 ENE 4018 0.268E-05 0.392E-06 E 4018 0.254E-05 0.391E-06 ESE 4018 0.277E-05 0.390E-06 SE 4018 0.262E-05 0.327E-06 SSE 4018 0.239E-05 0.267E-06 S 4018 0.246E-05 0.317E-06 All Direction Case 0.330E-05 0.490E-06

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-76 REV. 11, JANUARY 2005 TABLE 2.3-42 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 624 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 975 0.488E-05 0.159E-05 SW 975 0.670E-05 0.229E-05 WSW 975 0.643E-05 0.244E-05 W 975 0.711E-05 0.258E-05 WNW 975 0.584E-05 0.235E-05 NW 975 0.746E-05 0.312E-05 NNW 975 0.888E-05 0.322E-05 N 975 0.984E-05 0.402E-05 NNE 975 0.886E-05 0.401E-05 NE 975 0.750E-05 0.351E-05 ENE 975 0.706E-05 0.229E-05 E 975 0.654E-05 0.287E-05 ESE 975 0.826E-05 0.275E-05 SE 975 0.568E-05 0.215E-05 SSE 975 0.493E-05 0.152E-05 S 975 0.551E-05 0.153E-05 All Direction Case 0.810E-05 0.296E-05

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-77 REV. 11, JANUARY 2005 TABLE 2.3-43 CLINTON POWER STATION SITE ACCIDENT /Q CALCULATIONS 624 HOUR AVERAGING PERIOD (SEC. PER CUBIC METER)

DOWNWIND DISTANCE 5 PCT 50 PCT SECTOR (METERS) /Q /Q SSW 4018 0.101E-05 0.270E-06 SW 4018 0.138E-05 0.382E-06 WSW 4018 0.120E-05 0.402E-06 W 4018 0.149E-05 0.435E-06 WNW 4018 0.114E-05 0.391E-06 NW 4018 0.145E-05 0.533E-06 NNW 4018 0.167E-05 0.552E-06 N 4018 0.178E-05 0.661E-06 NNE 4018 0.155E-05 0.664E-06 NE 4018 0.149E-05 0.605E-06 ENE 4018 0.139E-05 0.386E-06 E 4018 0.122E-05 0.491E-06 ESE 4018 0.153E-05 0.422E-06 SE 4018 0.104E-05 0.333E-06 SSE 4018 0.926E-06 0.231E-06 S 4018 0.103E-05 0.246E-06 All Direction Case 0.155E-05 0.480E-06

Period of Record - May 1972-April 1977.

CPS/USAR CHAPTER 02 2.3-78 REV. 11, JANUARY 2005 TABLE 2.3-44 ANNUAL AVERAGE /Q CLINTON: 5/72 THROUGH 4/77 MEAN ANNUAL /Q BY SECTOR; GROUND-LEVEL RELEASE SECTOR AVERAGE DOWNWIND DISTANCE (miles)

DOWNWIND SECTOR 0.5 1.5 2.5 3.5 SSW 8.30E-06 1.20E-06 5.00E-07 2.70E-07 SW 1.00E-05 1.50E-06 6.30E-07 3.40E-07 WSW 1.10E-05 1.60E-06 6.40E-07 3.50E-07 W 8.40E-06 1.40E-06 6.00E-07 3.40E-07 WNW 9.30E-06 1.50E-06 6.20E-07 3.50E-07 NW 9.10E-06 1.50E-06 6.30E-07 3.70E-07 NNW 1.00E-05 1.70E-06 7.20E-07 4.10E-07 N 1.48E-05 2.30E-06 1.03E-06 5.70E-07 NNE 1.80E-05 3.00E-06 1.30E-06 7.00E-07 NE 1.60E-05 2.50E-06 1.00E-06 5.60E-07 ENE 1.00E-05 1.60E-06 6.80E-07 3.90E-07 E 9.00E-06 1.40E-06 5.70E-07 3.30E-07 ESE 6.80E-06 1.20E-06 5.20E-07 3.00E-07 SE 8.30E-06 1.30E-06 5.20E-07 2.90E-07 SSE 7.40E-06 1.20E-06 5.30E-07 2.90E-07 S 6.20E-06 1.00E-06 4.10E-07 2.30E-07 ALL 1.63E-04 2.59E-05 1.09E-05 6.09E-06 44,232 HRS EXAMINED DOWNWIND SECTOR 4.5 5.0 7.5 10.0 SSW 1.80E-07 1.40E-07 7.10E-08 4.20E-08 SW 2.20E-07 1.80E-07 8.80E-08 5.40E-08 WSW 2.50E-07 1.90E-07 9.20E-08 5.60E-08 W 2.30E-07 1.90E-07 9.90E-08 6.20E-08 WNW 2.30E-07 1.90E-07 9.50E-08 5.80E-08 NW 2.40E-07 2.00E-07 1.00E-07 6.50E-08 NNW 2.70E-07 2.20E-07 1.10E-07 6.90E-08 N 3.80E-07 3.20E-07 1.65E-07 1.04E-07 NNE 4.70E-07 4.10E-07 2.10E-07 1.30E-07 NE 3.70E-07 3.10E-07 1.50E-07 9.30E-08 CPS/USAR CHAPTER 02 2.3-79 REV. 11, JANUARY 2005 TABLE 2.3-44 (CONT'D)

ANNUAL AVERAGE /Q CLINTON: 5/72 THROUGH 4/77 MEAN ANNUAL /Q BY SECTOR; GROUND-LEVEL RELEASE SECTOR AVERAGE DOWNWIND DISTANCE (miles)

ENE 2.50E-07 2.10E-07 1.00E-07 6.80E-08 E 2.20E-07 1.70E-07 8.80E-08 5.30E-08 ESE 1.90E-07 1.70E-07 8.70E-08 5.50E-08 SE 1.90E-07 1.60E-07 7.80E-08 4.70E-08 SSE 1.90E-07 1.60E-07 8.30E-08 5.00E-08 S 1.50E-07 1.30E-07 6.50E-08 4.00E-08 ALL 4.03E-06 3.35E-06 1.68E-06 1.05E-06 44,232 HRS EXAMINED DOWNWIND SECTOR 15.0 25.0 35.0 45.0 SSW 2.10E-08 8.50E-09 4.70E-09 3.00E-09 SW 2.60E-08 1.00E-08 5.70E-09 3.70E-09 WSW 2.70E-08 1.10E-08 6.00E-09 3.90E-09 W 3.10E-08 1.30E-08 7.60E-09 5.00E-09 WNW 2.90E-08 1.20E-08 7.20E-09 4.60E-09 NW 3.30E-08 1.40E-08 8.00E-09 5.40E-09 NNW 3.50E-08 1.50E-08 8.50E-09 5.60E-09 N 5.20E-08 2.30E-08 1.30E-08 8.50E-09 NNE 6.50E-08 2.80E-08 1.60E-08 1.00E-08 NE 4.60E-08 1.90E-08 1.10E-08 6.90E-09 ENE 3.30E-08 1.40E-08 8.20E-09 5.50E-09 E 2.70E-08 1.10E-08 6.30E-09 4.10E-09 ESE 2.80E-08 1.20E-08 7.30E-09 4.70E-09 SE 2.30E-08 9.60E-09 5.20E-09 3.50E-09 SSE 2.60E-08 1.10E-08 6.30E-09 4.00E-09 S 2.00E-08 8.50E-09 4.90E-09 3.20E-09 ALL 5.22E-07 2.20E-07 1.26E-07 8.16E-08 44,232 HRS EXAMINED CPS/USAR CHAPTER 02 2.3-80 REV. 12, JANUARY 2007 TABLE 2.3-45 CLINTON POWER STATION JOINT WIND-STABILITY CLASS OCCURRENCE FREQUENCY DISTRIBUTION (2000-2002) 10 METER TOWER LEVEL Wind Direction Category Wind Speed Category(1)

N NNE NE ENE E ESESE SSES SSWSW WSWW WNW NW NNWTotal 2 0 1 2 6 4 16 13 19 6 5 7 5 3 2 2 1 92 3 18 7 62 63 76 89 94 105 13489 99 41 41 45 40 25 1028 4 41 37 97 33 25 14 50 32 140132129 71 89 80 91 43 1104 5 19 14 23 5 4 0 9 26 71 65 64 42 78 77 53 22 572 6 0 0 1 0 0 0 1 2 9 5 2 5 19 18 9 5 76 1 (A) 7 0 0 0 0 0 0 0 1 0 0 0 2 3 0 2 0 8 2 4 0 5 6 7 13 15 7 16 9 9 6 3 4 1 0 105 3 21 22 48 32 24 20 43 39 54 54 51 32 36 39 41 21 577 4 42 33 30 19 10 1 13 33 50 63 86 60 54 54 59 29 636 5 13 8 12 2 3 1 4 19 32 42 24 34 29 25 16 7 271 6 1 1 2 0 0 0 0 2 8 3 3 4 9 6 4 3 46 2 (B) 7 0 0 0 0 0 0 0 0 1 0 0 1 4 0 1 0 7 2 2 3 6 5 6 9 10 10 9 5 10 6 1 8 8 3 101 3 30 18 51 35 23 23 37 32 36 27 40 29 32 37 44 40 534 4 41 41 29 18 6 8 19 33 64 50 58 45 49 54 42 29 586 5 17 21 22 0 1 0 8 12 25 20 19 27 31 52 33 13 301 6 0 9 5 0 0 0 0 1 6 4 2 16 9 12 7 1 72 3 (C) 7 0 0 0 0 0 0 0 0 2 0 0 1 1 1 0 0 5 2 19 23 32 38 53 70 35 38 22 33 42 22 32 23 24 21 527 3 153 160 231 183 157 180210230 173181161 102170 159 177 1462773 4 232 219 226 106 51 72 147275 365360183 172312 315 296 1873518 5 60 79 73 7 1 2 27 71 18321060 93 215 193 130 53 1457 6 2 9 8 0 0 0 0 5 47 30 9 24 42 31 10 4 221 4 (D) 7 1 1 1 0 0 0 0 0 0 0 1 1 9 4 4 1 23 2 18 36 67 65 87 10989 67 63 68 55 55 37 31 30 17 894 3 77 96 197 148 142 152214361 407314198 177154 161 135 1023035 4 42 43 33 21 11 13 67 178 350347160 108124 93 38 46 1674 5 0 6 1 0 0 0 5 17 11276 33 24 42 10 10 5 341 6 0 0 0 0 0 0 0 3 18 6 1 1 1 2 1 0 33 5 (E) 7 0 0 3 0 0 0 0 1 0 0 0 0 0 0 0 0 4 2 21 50 90 63 60 64 51 48 52 63 44 53 46 40 30 16 791 3 38 83 136 72 37 21 89 98 10211199 84 73 51 78 23 1195 4 1 8 10 16 11 1 3 11 23 24 17 35 9 15 14 6 204 5 1 0 0 1 0 0 0 0 0 0 4 13 2 0 1 3 25 6 0 0 0 0 0 0 0 0 0 0 0 3 0 0 0 0 3 6 (F) 7 0 1 3 0 0 0 0 0 1 0 0 0 0 0 0 0 5 2 20 74 127 50 45 35 27 25 21 24 30 39 38 50 42 16 663 3 10 65 102 21 17 4 21 15 21 19 31 28 16 25 53 4 452 4 0 1 1 5 5 0 0 0 0 0 4 4 0 0 7 1 28 5 0 0 0 2 4 0 0 0 0 0 4 4 0 0 0 0 14 6 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 1 7 (G) 7 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 Notes: 1) Wind speed categories defined as follows:

Category Wind Speed (mph) 2 >=0.5 to <3.5 3 >=3.5 to <7.5 4 >=7.5 to <12.5 5 >=12.5 to <18.5 6 >=18.5 to <24 7 >=24 2) Wind speed category 1 is assumed for calms. Calm occurrences by stability cla ss: A=0, B=2, C=1, D=3, E=7, F=10, G=10

CPS/USAR CHAPTER 02 2.3-81 REV. 12, JANUARY 2007 TABLE 2.3-46 CLINTON POWER STATION JOINT WIND-STABILITY CLASS OCCURRENCE FREQUENCY DISTRIBUTION (2000-2002) 60 METER TOWER LEVEL Wind Direction Category Wind Speed Category(1)

N NNE NE ENE E ESESE SSES SSWSWWSWW WNW NW NNWTotal 2 0 1 2 2 2 20 20 15 7 5 3 1 2 1 1 2 84 3 8 10 28 31 50 65 73 76 63 54 53 35 25 19 18 14 622 4 29 43 47 49 54 23 47 69 95 12811251 57 46 79 47 976 5 36 17 71 22 19 4 23 27 85 11284 58 90 47 57 37 789 6 1 5 18 6 5 0 3 22 46 35 26 11 33 36 35 17 299 1 (A) 7 0 9 7 2 2 0 2 15 12 6 7 7 23 14 8 9 123 2 2 0 0 2 7 9 16 15 6 13 7 5 3 3 2 1 91 3 11 20 18 18 17 28 23 19 31 32 34 19 21 26 23 17 357 4 35 22 29 15 13 10 16 33 39 49 75 50 49 29 30 53 547 5 27 14 18 13 10 0 8 20 42 65 50 38 39 30 32 16 422 6 2 4 15 3 1 0 1 14 17 24 9 13 14 14 11 8 150 2 (B) 7 4 5 6 0 1 2 1 7 13 4 2 9 15 6 4 3 82 2 0 3 2 2 5 5 10 5 1 5 6 3 3 6 8 3 67 3 21 17 20 15 17 23 20 28 22 16 26 19 21 14 31 21 331 4 51 24 21 9 16 9 21 28 30 41 52 36 42 24 40 35 479 5 33 22 16 13 4 2 14 20 42 46 44 23 29 40 27 32 407 6 5 17 12 7 2 0 4 9 16 12 14 20 26 18 26 10 198 3 (C) 7 2 10 17 0 1 0 3 4 11 9 1 19 15 8 15 7 122 2 11 11 15 8 16 21 16 12 11 13 10 12 18 9 13 10 206 3 94 69 58 38 63 10770 66 56 79 70 44 68 69 92 89 1132 4 215 129 161 98 10486 168175195177159114158 140 187 1912457 5 199 183 158 129 80 45 73 204327314168165224 229 284 1652947 6 38 62 104 41 24 3 34 10518416055 80 174 100 114 39 1317 4 (D) 7 11 18 38 9 2 4 11 36 98 66 14 37 74 21 33 14 486 2 12 9 11 3 5 11 10 6 12 5 8 13 7 4 6 5 127 3 36 24 27 27 39 11296 72 45 62 47 42 35 27 37 30 758 4 80 78 90 106 92 86 158257253187159140113 88 110 99 2096 5 70 91 107 73 82 15 65 244422380236146130 96 85 51 2293 6 1 11 10 5 17 1 17 59 14913863 31 25 9 6 6 548 5 (E) 7 9 10 4 1 2 8 2 10 73 15 7 9 10 2 1 2 165 2 7 6 2 0 13 12 0 5 9 4 6 4 5 5 7 3 88 3 26 8 21 21 25 46 35 26 31 28 35 14 12 19 14 15 376 4 38 33 52 55 50 29 43 58 91 69 59 61 46 45 36 60 825 5 18 72 102 72 49 2 13 31 83 83 11187 45 20 29 27 844 6 0 2 8 3 4 0 2 6 9 13 9 10 2 0 0 4 72 6 (F) 7 6 8 1 0 0 0 0 0 1 0 0 5 0 1 2 1 25 2 6 2 7 7 3 16 3 1 9 8 7 5 4 3 6 8 95 3 5 15 12 10 25 32 17 22 13 14 15 12 13 14 16 20 255 4 15 25 36 56 58 19 14 23 28 42 27 29 20 13 39 25 469 5 10 25 57 52 36 0 1 6 13 16 29 30 9 3 10 18 315 6 0 2 2 4 7 0 0 0 0 0 2 6 0 0 0 1 24 7 (G) 7 1 0 1 0 0 0 0 1 0 0 1 4 0 0 0 0 8 Notes: 1) Wind speed categories defined as follows:

Category Wind Speed (mph) 2 >=0.5 to <3.5 3 >=3.5 to <7.5 4 >=7.5 to <12.5 5 >=12.5 to <18.5 6 >=18.5 to <24 7 >=24 2) Wind speed category 1 is assumed for calms. Calm occurrences by stability cla ss: A=0, B=2, C=1, D=3, E=7, F=10, G=10 CPS/USAR CHAPTER 02 A2.3-1 REV. 12, JANUARY 2007 ATTACHMENT A2.3 ANALYTICAL FOG MODEL CPS/USAR CHAPTER 02 A2.3-1a REV. 11, JANUARY 2005 ATTACHMENT A2.3 - ANALYTICAL FOG MODEL The basic problem of predicting steam fog from a warm lake requires the following calculations: a. Determine the evaporation per unit area of the lake. b. Estimate the amount of evaporated water vapor that will be condensed due to existing ambient conditions. c. Calculate the expected downwind concentrations of condensed water vapor. d. Relate the calculated condensed water vapor to horizontal visibility. A model has been developed by TRC that calculates values for a, b, and c. A relationship between visibility and condensed water vapor has been established and is used to relate the computed values to expected visibility. The model has been calibrated by means of conditions recorded at a large cooling pond for a nuclear power station (Reference 1). The observed data included water temperatures at various parts of the lake, ambient air temperature and relative humidity, and observations of ambient fog and lake steaming by trained weather observers. For the Clinton Power Station (CPS) study, only air temperatures above -4x F were used. This allows most of the condensed water vapor to form water droplets rather than condense directly into the solid phase as ice crystals. Evaporation from a unit of surface on the lake per unit of time in a layer from h 1 to h2 is computed by Equation A2.3-1: ()()()2121221h/hn1uuqq2E= (cgs units) (A2.3-1) Where k is the Von Karmen coefficient: r is the density of air; h2 and h1 are the heights of the top and bottom, respectively, of the layer in which evaporation takes place; q 1 and q2 are the specific humidities: and u 1 and u2 are the wind speeds (References 2 and 3). The value of E from this equation is converted to an equivalent line source value using the dimensions of the unit area. The proper values of wind fetch for use in defining a unit area for conversion into a line source were examined and evaluated in the calibration and verification of the model. The standard line source diffusion equation for surface concentration is as follows (Reference 4):

()()()()()yoyoZ2yyerf2yyerf2Eo,y,xX+= (A2.3-2) The equation is used to calculate concentrations of water vapor and condensed water vapor for an orthogonal array of points downwind. In Equation A2.3-2, E is the line source strength determined from Equation A2.3-1, erf is the error function, y o is the half-width of the source, and y and z are the horizontal and vertical diffusion parameters, respectively. The predicted CPS/USAR CHAPTER 02 A2.3-2 REV. 11, JANUARY 2005 concentrations are used to determine the amount of condensed water vapor that would exist at the downwind points, after allowing the ambient air to reach saturation. The downwind concentrations are related to horizontal visibility by the following relationship:

V____rkcw= (A2.3-3) where w is the liquid water content of the air and V is the resulting visibility. The term c is an empirical constant and k is a factor that accounts for drop size distribution around the average radius r. This equation was derived in other research studies on fog where both liquid water content and the drop size distribution were measured. However, such studies on natural fog (Reference 5) include a predominate number of warm fog cases in which the drop size distribution is different than for cold fogs. Therefore, the data for natural fog are used when the ambient air temperature is 36

° F or higher. For cold fogs, a mean drop size radius of 10m was used with a factor of k = 1.2 in Equation A2.3-1 (Reference 1). This produces a curve that is used when the air temperature is 28

° F or less, and is in good agreement with the results of a U.S. Army study on arctic fogs (Reference 6). A log-log plot of Equation A2.3-3 is presented in Figure A2.3-1 for the warm fog and cold fog cases. An interpolation is used between the two curves for transition temperatures between 28

° F and 36

° F. Occurrences of overpredicting downwind concentrations of water vapor were investigated as part of the model development. The problem was related to the evaporative processes on a parcel of air as it travels across the lake. That is, the term (q 1 - q2) from Equation A2.3-1 decreases with travel time because of the following dynamic effects; a. The specific humidity of the air q 2, initially is a function of the dew point, and is normally less than the saturation specific humidity. As the air receives water vapor from the pond, saturation is reached, increasing the value of q

2. b. As further moisture is received by the air after it has reached saturation, the water vapor condenses into liquid water, releasing the latent heat of condensation of the water vapor. This further increases q

2. c. As fog is formed, heat radiated from the pond is reflected and absorbed by the water droplets, further increasing the air temperature and hence, q

2. d. Convection of heat from the pond surface to the atmosphere still further increases q

2. As the value of q 2 increases by the previous methods, the term (q 1 - q2) decreases and hence the evaporation into a parcel of air decreases as it travels across the lake. The first two mechanisms are quantifiable and were used to determine the weighting factor for adjusting the evaporation rate with travel time. Radiation and convective effects were not computed and thus were empirically accounted for in the calibration of the model.

CPS/USAR CHAPTER 02 A2.3-3 REV. 11, JANUARY 2005 Use of the Model Predicted water temperatures for six areas of the lake evaluated to date are reduced to representative (monthly) values. The lake is divided into adjoining rectangular blocks that present an edge perpendicular to the wind direction to be evaluated. Each of these blocks is used as a source area to compute the evaporation-condensation-diffusion process over the lake and surrounding areas of interest. To evaluate the potential for steam fog and subsequent drift off the lake, an ambient air temperature, relative humidity, wind direction, wind speed, and atmospheric stability are input to the model for a given lake source area and water temperature. The model output is water vapor concentration at orthogonal grid points that cover the area of interest. A grid mesh of 500 meters is normally used, but was frequently varied to determine the location of critical values of water vapor concentration.

CPS/USAR CHAPTER 02 A2.3-4 REV. 11, JANUARY 2005 A2.3 References

1. R. R. Hippler, "Analytical Study of Cooling Pond Fogging at Low Temperatures for Commonwealth Edison Company," The Research Corporation of New England, Wethersfield, Connecticut, 1972. 2. C. W. Thornthwaite and B. Holzman, "The Determination of Evaporation from Land and Water Surfaces," Monthly Weather Review, January 1939. 3. T. F. Malone (Ed.), "Compendium of Meteorology," American Meteorological Society, Boston, Massachusetts, P. 505, 1951. 4. D. B. Turner, "Workbook of Atmospheric Dispersion Estimates," U.S. Dept. of Health, Education and Welfare, Public Health Service Publication No. 999-AP-26, p. 40, 1969. 5. T. F. Malone (Ed.), "Compendium of Me teorology," American Meteorology Society, Boston, Massachusetts, p. 1180, 1951. 6. M, Kumai, "Arctic Fog Droplet Size Distribution and Its Effects on Light Attenuation," U.S. Army Cold Regions Research and Engineering Laboratory, Hanover, New Hampshire,

April 1972.

CPS/USAR CHAPTER 02 2.5-1 REV. 11, JANUARY 2005 2.5 GEOLOGY, SEISMOLOGY, AND GEOTECHNICAL ENGINEERING The Clinton Power Station site is located in the central stable region of North America in the Illinois Basin, slightly west of the La Salle Anticlinal Belt, approximately 6 miles east of the city of Clinton, DeWitt County, Illinois. The site area consists of a gently rolling upland developed on ground moraine, which has been dissected by the southwest-flowing Salt Creek and the North Fork of Salt Creek. Topographic relief varies from approximately 10 feet on the upland to a maximum of about 80 feet between the upland and valley bottoms. Strata underlying the site consists of an estimated 170 to 360 feet of Quaternary overburden, largely Wisconsinan, Illinoian, and pre-Illinoian aged glacial deposits resting on essentially flat-lying Pennsylvanian-aged shales, sandstones, and thin coal beds. A horizontal ground surface acceleration of 25% of gravity (0.25g) applied at foundation level was selected for the safe shutdown earthquake (SSE). This acceleration value was derived by assuming an Intensity VIII event (MM) centered near the site. Extensive geotechnical investigations carried out prior to and during construction (including geologic mapping of the excavations) showed nothing that would preclude safe construction or operation of a nuclear-fueled power station. There are no known faults or folds of design significance at or anywhere near the site. Major power block structures were constructed on mat foundations underlain by compacted fill resting on hard Illinoian till. The cooling lake was formed by construction of an earth-filled dam across Salt Creek downstream of its confluence with the North Fork of Salt Creek. The emergency cooling ultimate heat sink was formed by constructing a submerged pond within the cooling lake reservoir in the North Fork of Salt Creek. Required pond capacity was developed by construction of an earth dam across the natural stream channel and excavation upstream from the berm. Adequate borrow was available for all fills. Suitable founding strata were available for all structures. Geologic investigations undertaken include: a review of published and unpublished data; discussions with individuals, agencies, and companies having information in the region or site area; reconnaissance field investigations, drilling and sampling, and surface and borehole geophysics; laboratory testing of soil, rock, and water samples, and detailed geologic mapping of excavations. The firms who performed the work are as follows: INVESTIGATIONS PERFORMED BY Geologic Literature Review Dames & Moore, and Sargent & Lundy

Geologic Reconnaissance Dames & Moore

Geologic Mapping of Excavations Sargent & Lundy

Test Borings Dames & Moore, and Sargent & Lundy

Geophysical Explorations Dames & Moore

CPS/USAR CHAPTER 02 2.5-2 REV. 11, JANUARY 2005 Laboratory Tests Dames & Moore, U.S. Testing, Westenhoff &

Novick, and Soil Testing Services Inc.

Foundation Considerations

Subsurface Conditions Dames & Moore, and Sargent & Lundy

Foundations Sargent & Lundy

Vibratory Ground Motion Dames & Moore, and Sargent & Lundy

Surface Faulting Dames & Moore, and Sargent & Lundy

Stability of Subsurface Materials Dames & Moore, Sargent & Lundy, and

Woodward-Clyde

Consultants Stability of Slopes Sargent & Lundy

Embankments and Dams Sargent & Lundy

Groundwater Dames & Moore, and Sargent & Lundy

2.5.1 Basic Geologic and Seismic Information 2.5.1.1 Regional Geology The region surrounding the site lies within the Central Stable Region of the North American Continent (Reference 1). This province is a tectonically stable area characterized by gently dipping sedimentary rock of Paleozoic overlain by thin Cenozoic deposits mostly quarternary glacial drift, and, locally by Mesozoic strata. Beneath the Paleozoic is a basement complex of Precambrian and igneous and metamorphic rocks. Intermittent slow subsidence and gentle uplift through the Paleozoic has resulted in broad basins (e.g., the Illinois, Michigan, and Forest City Basins), filled with gently dipping sedimentary rocks, and in intervening broad arches or highs (e.g., the Kankakee Arch, Mississippi River Arch, etc.). Locally, folds and faults have been superimposed on this pattern. The Clinton site is located on the northwest flank of the Illinois Basin, west of the La Salle Anticlinal Belt. The Paleozoic sedimentary rock sequence is punctuated by several unconformities of regional importance, reflecting widespread advances and withdrawals of the Paleozoic seas across the interior of North America.

CPS/USAR CHAPTER 02 2.5-3 REV. 11, JANUARY 2005 Locally, in the regional area, there was Cretaceous and Tertiary deposition in a few local areas in the Mississippi Embayment, western Illinois, eastern Missouri, and southern Indiana. During Quaternary time, widespread deposition occurred in the regional area as the result of continental glaciation. Approximately 170 to 360 feet of Quaternary deposits overlie the Pennsylvanian bedrock at the Clinton site. These deposits consist of a complex sequence of glacial drift, stream alluvium, and loess. 2.5.1.1.1 Regional Physiography The region of the United States in which the Clinton Power Station site is located is part of the Till Plains Section of the Central Lowland Physiographic Province (Reference 2). The terrain aspect of central Illinois and adjacent Indiana is typical of the province, consisting of undulating, low-relief topography formed by a glacial drift cover whose thickness ranges from a few tens of feet to several hundreds of feet. Much of the Till Plains Section is characterized by landforms of low, commonly arcuate ridges, called moraines, interspersed with relatively flat intermorainal areas. The Clinton Power Station site is situated in a sector of the Till Plains Section known as the Bloomington Ridged Plain as shown in Figure 2.5-1. Postglacial stream development has dissected the drift mantle and in some areas along the main valleys, preglacial bedrock has been exposed by erosion; however, there are no bedrock exposures near the site area. Elevations on the general drift surface between drainageways in the general area of the site average about 740 feet above sea level. 2.5.1.1.2 Regional Stratigraphy Overburden deposits consisting of Quaternary-aged glacial drift and stream alluvium overlie thick sequences of Paleozoic sedimentary rock throughout most of Illinois and adjacent Indiana. In the extreme northern part of Illinois, the drift rests principally upon Ordovician and Silurian formations. Elsewhere, the uppermost strata beneath the glacial drift consist mainly of Pennsylvanian-aged (Late Paleozoic) rocks. Figure 2.5-2 illustrates the general rock sequence in central Illinois; and Figure 2.5-3 shows the regional bedrock geology of Illinois and surrounding states. Most of the Paleozoic formations in Illinois dip gently (about 25 feet per mile) with some thickening toward the axis of the Illinois Basin in southeastern Illinois. Figure 2.5-4 shows the regional stratigraphic relationships along north-south and east-west cross sections near the site. The ages for the geologic periods discussed below are taken from Faul

(Reference 3). 2.5.1.1.2.1 Cenozoic Era (Present to 65 +/- 2 million years B.P.)

2.5.1.1.2.1.1 Quaternary System (Present to 2 +/- 1 million years B.P.)

The surficial deposits of most of the regional area are Quaternary in age and are classified as part of the Pleistocene Series. The deposits consist predominantly of glacial or glacially-derived sediments of glacial till, outwash, loess (a wind-blown silt), and glaciolacustrine deposits, as well as alluvium. The age and geomorphic relations of the glacial deposits in northern and central Illinois are shown in Figure 2.5-1. The stratigraphic sequence of Pleistocene deposits in Illinois is shown in Figure 2.5-5.

CPS/USAR CHAPTER 02 2.5-4 REV. 11, JANUARY 2005 There were four major periods of glaciation during Pleistocene time in the regional area. From oldest to youngest, these periods are known as the Nebraskan, Kansan, Illinoian, and Wisconsinan Stages. (The present classification of pre-Illinoian Drift, however, has been questioned by some, see Reference 4.) During each of these glacial periods, glaciers periodically advanced and retreated across parts of the regional area. Consequently, a complex sequence of deposits developed, as evidenced by the stratigraphic classification for the Pleistocene deposits of Illinois (Figure 2.5-5). Nebraskan and Kansan age glacial deposits are present at the surface and in the subsurface of the regional area in Iowa, Missouri, and parts of western and east-central Illinois. Illinoian age deposits are present beyond the limit of Wisconsinan glaciation in northern and central Illinois (Figure 2.5-1). Illinoian age deposits are also found beneath the Wisconsinan drift cover up to 20 to 40 miles back from the Wisconsinan front (Reference 2). The site is located a few miles inside the limit of Wisconsinan glaciation (Figure 2.5-1). Quaternary deposits in the regional area locally exceed 400 feet in thickness but are generally much thinner. They cover an irregular bedrock surface, largely erosional in origin characterized by valleys and uplands that were developed before and during glacial time (Reference 6). The locations of buried bedrock valleys in Illinois are shown in Figures 2.5-6 and 2.5-7. The sequence of Pleistocene deposits in the regional area is comprised of glacial or glacially-derived sediments of glacial till, outwash, loess, and glaciolacustrine deposits. Glacial till is a poorly sorted sediment consisting of a matrix of sand, silt, and clay with interspersed pebbles or cobbles. Individual glacial till units tend to be texturally and mineralogically distinctive and uniform, and can be traced over wide areas (References 5 and 7). Glacial till can be deposited by a number of subglacial and supraglacial processes. Discontinuous lenses of stratified sand, silt, or gravel may occur within glacial till and are associated with melt waters from glacial ice

(Reference 8). Outwash deposits consist dominantly of sand and gravel with minor amounts of silt and clay. The texture of outwash deposits can vary greatly both laterally and vertically due to differences in the characteristics of the meltwater channels, and differences in meltwater velocity caused by diurnal and seasonal differences in the rate of meltwater discharge. Loess is a wind-blown silt derived principally from the fine-grained sediments along meltwater rivers and channels. Loess thickness decreases eastward away from the major rivers in the regional area which served as major drainageways for meltwater streams. In the bluffs bordering the eastern side of major river valleys, loess typically is on the order of a few tens of feet in thickness thinning to only a few feet in thickness a short distance away from the river bluffs. Glaciolacustrine deposits were formed in lakes, on or fed by glaciers. Glaciolacustrine deposits are predominantly silts and clays, although they may locally be sands or gravels near the margins where they were parts of beaches, bars, or deltas. Between glacial periods, the climate returned to more temperate conditions. As the glacial and glacially-derived sediments were exposed, weathering processes began to alter them, and soils were formed. The thickness and character of the resulting soils are largely a function of climate, topographic position, vegetation, and duration of the interglacial stage.

CPS/USAR CHAPTER 02 2.5-5 REV. 11, JANUARY 2005 The youngest deposits in the regional area are classified as Holocene in age (Reference 2). They consist of alluvium formed along present rivers and drainageways, colluvium, dune sand, peat, and lacustrine deposits formed in present lakes. 2.5.1.1.2.1.2 Tertiary System (2 +/- 1 to 65 +/- 2 million years B.P.)

Late Tertiary deposits consist of chert gravels in the Mississippi Embaymment, of high-level, Plio-Pleistocene isolated patches of chert gravel in western Illinois (Reference 2) and eastern Missouri (Reference 9), and of Mio-Pliocene chert gravels in southern Indiana (Reference 10.) 2.5.1.1.2.2 Mesozoic Era (65 +/- 2 to 225 +/- 5 million years B.P.)

2.5.1.1.2.2.1 Cretaceous System (65 +/- 2 to 135 +/- 5 million years B.P.)

Cretaceous age sediments in the regional area are present only in a small area in western Illinois and in the Mississippi Embayment area of extreme southern Illinois (Figure 2.5-3), and are Late Cretaceous in age. In western Illinois the Cretaceous age rocks are up to 100 feet thick and consist of a basal gravel overlain by sands and clayey sands. These strata apparently are the easternmost outliers of Cretaceous sediments that formerly covered the region east of the Rocky Mountains and north of the Ozarks (Reference 2). In southern Illinois, the Cretaceous age rocks of the Mississippi Embayment area consist of up to 500 feet of chert gravels, deltaic sands, and marine clays (Reference 2). 2.5.1.1.2.2.2 Jurassic System (135 +/- 5 to 190 +/- 5 million years B.P.)

There are no known deposits of Jurassic age in the regional area.

2.5.1.1.2.2.3 Triassic System (190 +/- 5 to 225 +/- 5 million years B.P.)

There are no known deposits of Triassic age in the regional area.

2.5.1.1.2.3 Paleozoic Era (225 +/- 5 to approximately 600 million years B.P.)

2.5.1.1.2.3.1 Permian System (225 +/- 5 to 270 +/- 5 million years B.P.)

There are no known deposits of Permian age in the regional area.

2.5.1.1.2.3.2 Pennsylvanian System (270 +/- 5 to 320 +/- 10 million years B.P.)

The bedrock surface throughout much of Illinois has been developed on strata of Pennsylvanian age, as shown in Figure 2.5-3. This surface is covered by Pleistocene drift except where erosion has removed these glacial materials. The base of the Pennsylvanian is an unconformity of major regional extent. The Pennsylvanian sequence is composed of cyclothem units formed by alternating beds of shale, sandstone, coal, limestone, and siltstone. The thickness of the Pennsylvanian age rocks varies greatly depending upon the amount of subsequent erosion, but it is generally on the order of a few hundred feet to 700 or 800 feet in the site region. Further south, the Pennsylvanian in the Illinois Basin thickens to more than 2200 feet.

CPS/USAR CHAPTER 02 2.5-6 REV. 11, JANUARY 2005 2.5.1.1.2.3.3 Mississippian System (320 +/- 10 to 340 +/- 10 million years B.P.)

Strata of the Mississippian System outcrop along the outer margins of the Illinois Basin (Figure 2.5-3), and are present in the subsurface of the basin as well. Mississippian-age sediments consist predominantly of limestone, with lesser siltstone and shale, and are on the order of 500 to 600 feet in thickness near the site area. 2.5.1.1.2.3.4 Devonian System (340 +/- 10 to 400 +/- 10 million years B.P.)

Devonian age rocks, consisting predominantly of shale and limestone, underlie much of the general region. These sediments are on the order of about 200 feet thick near the area of the site. There is a major unconformity at the base of Middle-Devonian strata. 2.5.1.1.2.3.5 Silurian System (400 +/- 10 to 430 +/- 10 million years B.P.)

Silurian age rocks underlie the Devonian strata in the general region and outcrop over large areas of northern Illinois and Indiana beneath a cover of glacial drift. Near the site area Silurian strata form the bedrock surface in parts of the La Salle Anticlinal Belt. The Silurian rocks are dominantly carbonates, some of which include reef structures. Silurian rock in the site area is approximately 450 feet thick (Reference 2). 2.5.1.1.2.3.6 Ordovician System (430 +/- 10 to approximately 500 million years B.P.)

Ordovician age rocks outcrop in the northern, southwestern, and southeastern parts of the regional area (Figure 2.5-3), and are present in the subsurface of most of the rest of the regional area. The Ordovician has been divided into three series, which from oldest to youngest are the Canadian, Champlainian, and Cincinnatian Series. The Canadian Series (Lower Ordovician) is represented by dolomite and sandstone. A major erosional unconformity separates the top of the Canadian Series from the basal sandstone of the overlying Champlainian Series. Strata of the Champlainian Series include the basal sandstone (St. Peter Sandstone) and an overlying sequence of limestone and dolomite. Upper Ordovician rocks of the Cincinnatian Series are dominantly shales. Ordovician rocks in the site area are approximately 1500 feet thick (Reference 2). 2.5.1.1.2.3.7 Cambrian System (Approximately 500 to 600 million years B.P.)

The Cambrian in Illinois and adjoining states is represented by only the Upper Cambrian Croixan Series, a sequence of sedimentary rocks ranging from 1500 to over 3500 feet in thickness in Illinois (Reference 2). Approximately the lower two-thirds of the Croixan Series is sandstone in northern and central Illinois, grading to dolomite further south. The upper one-third of the Croixan Series consists of interbedded siltstone, shale, sandstone, and dolomite.

Cambrian strata rest on the Precambrian basement with profound unconformity. Approximately 3100 feet of Cambrian rocks underlie the site (Reference 2). 2.5.1.1.2.4 Precambrian Era (Greater than approximately 600 million years B.P.)

Precambrian igneous rocks, approximately 1.1 to 1.4 billion years in age lie below the surface throughout the regional area. Precambrian rocks are at the surface in the Ozarks, more than 185 miles southwest of the site (Figure 2.5-3), but are generally deeply buried in the regional area, ranging in depth from 2000 feet near the Illinois-Wisconsin border to over 13,000 feet in the deepest part of the Illinois Basin in southeastern Illinois. Precambrian rocks in the regional CPS/USAR CHAPTER 02 2.5-7 REV. 11, JANUARY 2005 area are dominantly granite with associated granodiorite, rhyolite, felsite, or granophyre of closely related composition (Reference 2). 2.5.1.1.3 Historical Geology The geologic history of the regional area is discussed by geologic eras, which are subdivided into periods. The strata formed during the periods are classified as time-rock units and are designated as systems. In the following discussions, the periods are divided into early, middle, and late; the corresponding strata are designated as series and referred to as lower, middle, and upper, respectively. Regional stratigraphy is discussed in Subsection 2.5.1.1.2, and regional structural geology is discussed in Subsection 2.5.1.1.4. The ages given represent the broad time spans and are not restricted to those portions of the time interval represented by the rocks within the regional area. 2.5.1.1.3.1 Precambrian Era (Greater than approximately 600 million years B.P.)

Except for a few isolated exposures in the Ozarks, the Precambrian basement complex within the regional area is covered by younger strata. Outcrop and subsurface data on the Precambrian rocks indicate that the Precambrian basement in the regional area consists of igneous rocks, chiefly granite and associated rocks, that formed approximately 1.4 to 1.1 billion and in part 0.64 billion years ago (Reference 2). By their ages, the Precambrian rocks are assigned to the Precambrian Y division of the U.S. Geological Survey (Reference 11). A period of erosion, lasting 600 to 900 million years (Reference 2), followed the formation of the Precambrian basement, and resulted in up to 2000 feet of relief on the Precambrian surface in the Ozarks (Reference 12), and several hundred feet of relief elsewhere in the regional area (Reference 2). 2.5.1.1.3.2 Paleozoic Era (Approximately 600 to 225 +/- 5 million years B.P.)

2.5.1.1.3.2.1 Cambrian Period (Approximately 600 to 500 million years B.P.)

Within the regional area, deposition of sediments on the Precambrian basement did not begin until Late Cambrian time when shallow seas transgressed from the south. Initial sediments in the regional area consisted of a sequence of sandstones (Reference 2). These basal Cambrian clastics (Potsdam Sandstone Megagroup) grade southward and upward into carbonates. 2.5.1.1.3.2.2 Ordovician Period (Approximately 500 to 430 +/- 10 million years B.P.)

Sedimentation in the regional area continued from Late Cambrian into Early Ordovician time with no appreciable break. Early Ordovician sediments in the regional area are dominantly limestones and dolomites with some sandstones (Reference 2). The Reelfoot Basin in western Kentucky, outside of the regional area, appears to have been the center of thickest deposition in Early Ordovician time (Reference 13). Regional uplift occurred in the interval between Early and Middle Ordovician time, resulting in a major unconformity between rocks of Early and Middle Ordovician age throughout much of the central interior region, and in the development of karst in Illinois (Reference 14). During this CPS/USAR CHAPTER 02 2.5-8 REV. 11, JANUARY 2005 interval, the Wisconsin Dome was uplifted, and the Kankakee Arch began to form, separating the Illinois Basin from the Michigan Basin (Reference 14). The Kankakee Arch at this time was at a position somewhat northeast of its later location (Reference 14). Uplift also occurred in the southwestern part of the regional area in the Ozarks. In Middle Ordovician time, shallow seas readvanced into the area. The initial deposits were sandstone, followed by an extensive sequence of carbonates. In the interval between Middle and Late Ordovician time, the shallow seas retreated and then readvanced back into the regional area, as evidenced by the unconformity separating rocks of Middle and Late Ordovician age. Late Ordovician rocks in the regional area are dominantly shales, possibly reflecting uplift in the Appalachian region during the Taconic orogeny (Reference 14). The center of thickest deposition in the regional area migrated north during Ordovician time from the Reelfoot Basin, south of the regional area, to the southern Illinois area of the Illinois Basin and the Fairfield Basin (Reference 13). 2.5.1.1.3.2.3 Silurian Period (430 +/- 10 to 400 +/- 10 million years B.P.)

Rocks of Silurian age in the regional area were deposited on an erosional unconformity of low relief (Reference 14), which indicates a period of sea withdrawal and readvance in the interval from the end of Ordovician sedimentation to the beginning of Silurian sedimentation. The lower Silurian Series are predominantly carbonates, whereas Middle and Late Silurian strata are characterized by limestones or dolomites containing organic reefs and bioherms. Initial upwarp of the Sangamon Arch (see Subsection 2.5.1.1.4.1.13) may have occurred in Late Silurian time (Reference 14). The Wisconsin Dome underwent a second period of significant uplift after deposition of Silurian strata (Reference 15). The Lincoln Anticline also began to develop in Silurian or Devonian time (Reference 16). 2.5.1.1.3.2.4 Devonian Period (400 +/- 10 to 340 +/- 10 million years B.P.)

Early Devonian sedimentation was restricted to deposition of limestones and cherts in the southern part of the Illinois Basin. A major unconformity occurs at the base of the Middle Devonian in the regional area. This extensive unconformity is the result of regional uplift with withdrawal of the seas. The Ozark area was again uplifted during this time (Reference 17). Shallow seas readvanced into the regional area, and early-Middle Devonian sediments in the Illinois Basin were primarily limestones. In the late-Middle and Late Devonian, organic shale (New Albany Shale Group), derived from the Acadian Highland to the east, were deposited in the Illinois Basin (Reference 14). 2.5.1.1.3.2.5 Mississippian Period (340 +/- 10 to 320 +/- 10 million years B.P.)

Sedimentation in the Illinois Basin was generally continuous from Devonian into Mississippian time. By Middle Mississippian time, sedimentation conditions in the regional area had changed from dominantly clastic deposition to the accumulation of limestones and dolomites. Upper Mississippian sedimentation consisted of alternating deposits of clastics and carbonates. The main source of the Upper Mississippian age clastic sediments appear to have been from the CPS/USAR CHAPTER 02 2.5-9 REV. 11, JANUARY 2005 northeast, although some sediment came from the northwest and from the Ozark region (Reference 2). The principal folding along the Cap au Gres Faulted Flexure was post-Middle Mississippian and pre-Pennsylvanian, although minor movements occurred before and after this period. Development of the Mississippi River Arch began in Mississippian time and continued into Pennsylvanian time (Reference 14). 2.5.1.1.3.2.6 Pennsylvanian Period (320 +/- 10 to 270 +/- 5 million years B.P.)

A major unconformity of regional dimension separates rocks of the Mississippian and Pennsylvanian Series in the regional area. Strata in the regional area were warped, faulted, and truncated by erosion at the close of Mississippian time. Valleys as much as 450 feet deep were cut into Mississippian age strata and were subsequently filled with Pennsylvanian sediments (Reference 2). During the period marked by the Mississippian-Pennsylvanian unconformity, the La Salle Anticlinal Belt began to develop. As the La Salle Anticlinal Belt rose, the locus of maximum deformation moved progressively southward from the La Salle area (Reference 14). Renewed uplift also occurred in the Ozark area at this time, causing widespread erosion (Reference 17). Pennsylvanian sediments in the regional area were deposited in a gently subsiding trough, the Illinois Basin, that was open toward the south until post-Pennsylvanian time, when it was closed by uplift of the Pascola Arch, south of the regional area (Reference 2). The depositional environments during Pennsylvanian time were generally different than those of earlier Paleozoic time, and resulted in delta plain, brackish water, and marine sediments. Much of the sedimentation occurred in large deltas on the gently subsiding basin. Ninety to ninety-five percent of the Pennsylvanian strata consist of clastic rocks (Reference 2). Pennsylvanian age strata are characterized by vertical changes in lithology, commonly abrupt. In Illinois, 500 or more distinguishable units of sandstone, siltstone, shale, limestone, coal, and underclay grouped into cyclotherms can be distinguished (Reference 2). Pennsylvanian deposits thin over the La Salle Anticlinal Belt indicating continued uplift along this structure during Pennsylvanian time. The last period of folding along the La Salle Anticlinal Belt took place in post-Pennsylvanian time (Reference 18). Renewed uplift occurred in the Ozark area in post-Pennsylvanian time. This was followed by erosion over the entire state of Missouri (Reference 17). 2.5.1.1.3.2.7 Permian Period (270 +/- 5 to 225 +/- 5 million years B.P.)

No deposits of Permian age are present in the regional area. Permian sediments may have been deposited in the regional area (References 2 and 14), but if so, they have subsequently been removed by erosion. 2.5.1.1.3.3 Mesozoic Era (225 +/- 5 to 65 +/- 2 million years B.P.)

2.5.1.1.3.3.1 Triassic Period (225 +/- 5 to 190 +/- 5 million years B.P.)

There are no deposits of Triassic age in the regional area. This was largely a period of erosion in the regional area (Reference 2).

CPS/USAR CHAPTER 02 2.5-10 REV. 11, JANUARY 2005 2.5.1.1.3.3.2 Jurassic Period (190 +/- 5 to 135 +/- 5 million years B.P.)

There are no deposits of Jurassic age in the regional area. This was largely a period of erosion in the regional area (Reference 2). 2.5.1.1.3.3.3 Cretaceous Period (135 +/- 5 to 65 +/- 2 million years B.P.)

Beginning in Late Cretaceous time, shallow seas once again advanced into parts of the regional area, leaving beach, nearshore, or deltaic deposits in portions of southern and western Illinois (Figure 2.5-3). The Late Cretaceous age rocks in western Illinois are apparently the easternmost outliers of Cretaceous sediments that formerly covered the region east of the Rocky Mountains and north of the Ozarks. The sediments are characteristic of beach and nearshore deposits (Reference 2). The Late Cretaceous age rocks of extreme southern Illinois were part of a large delta formed where a river from the east discharged into the Mississippi Embayment (Reference 2). 2.5.1.1.3.4 Cenozoic Era (65 +/- 2 million years B.P. to the present) 2.5.1.1.3.4.1 Tertiary Period (65 +/- 2 to 2 +/- 1 million years B.P.)

There are no deposits of the Paleocene, Eocene, Oligocene, or Miocene epochs present in the regional area except for gravels in Indiana of possible Mio-Pliocene age (Reference 10). This was likely a period of erosion. Late Tertiary gravels in part of the Pliocene Series are present in the regional area in widely scattered outcrops (Reference 2). The Pliocene was primarily a time of erosion with some local areas of fluvial deposition. This deposition is represented by relict patches of chert and quartz gravel, part of which may be reworked from older Tertiary or Cretaceous gravels. Reworking of these Pliocene gravels may have continued into early Pleistocene (Reference 2). 2.5.1.1.3.4.2 Quaternary Period (2 +/- 1 million years B.P. to the present)

Most of the regional area was covered by continental ice sheets during Pleistocene time. Prior to the onset of Pleistocene glaciation, Tertiary erosion cycles had left the Paleozoic sediments as an essentially planar surface dissected by stream valleys (Reference 2). Advances of the continental ice sheets during substages within the Nebraskan (oldest), Kansan, Illionian, and Wisconsinian (youngest) stages have left a complex sequence of deposits. The glacial stages were separated by interglacial periods. Soils formed during these periods are generally well known and widely utilized stratigraphic horizons. Glaciations also altered the topography throughout much of the region and rearranged patterns (including the Mississippi River). Since the disappearance of the last of the Pleistocene glaciers, the region has been largely undergoing a period of erosion and most likely isostatic rebound. The youngest deposits in the regional area are classified as Holocene in age (Reference 2). They consist of alluvium formed along present rivers and drainageways, colluvium, dune sand, peat, and lacustrine deposits formed in present lakes.

CPS/USAR CHAPTER 02 2.5-11 REV. 11, JANUARY 2005 2.5.1.1.4 Regional Structural Geology The dominant bedrock structural features within the study region are illustrated in Figures 2.5-8, 2.5-9 and 2.5-10. An attempt has been made to show as many structural features within the study region as possible; however, some generalization has been necessitated by the small-scale mapping of closely spaced features such as zones of intense faulting in southern Illinois. Data tabulations for all important structures within a 200-mile radius are given in Tables 2.5-1 and 2.5-2 (see also Attachment D2.5). The dominant structures of the regional area and vicinity are the Illinois Basin and its bounding structures: the Mississippi River Arch, Wisconsin Arch, Kankakee Arch, Cincinnati Arch, Mississippi Embayment, Pascola Arch, Ste. Genevieve Fault Zone, and Lincoln Anticline. Within or crossing the Illinois Basin are major structures; such as the Sandwich Fault Zone, La Salle Anticlinal Belt, Rough Creek Fault Zone, and Wabash Valley Fault System; and structures of relatively less significance, such as the DuQuoin Monocline and Louden Anticline, for example. On the bounding arches are structures such as the Royal Center Fault and Fortville Fault. The movement on structures in the regional area was intermittent, and confined essentially to the Paleozoic. 2.5.1.1.4.1 Folding The site occupies the sediment-covered part of the craton near the interior part of the Illinois Basin and lies some 15 to 20 miles west of the La Salle Anticlinal Belt (see Figure 2.5-8). The distribution of major folds in the region is shown in Figure 2.5-9 and their characteristics are presented in Table 2.5-1. The knowledge of these structural features is based on surface and/or subsurface geological data. Some of the structures in the regional area have experienced post-Pennsylvanian movement. Due to the absence of sediments representing the interval from Pennsylvanian to Cretaceous or Pleistocene time, the age of final movement on these structures cannot be precisely dated. However, based on stratigraphic relationships and geologic history outside of the regional area, it is presumed that most of the post-Pennsylvanian tectonic activity is related to Appalachian structural development that occurred near the close of the Paleozoic Era (Reference 14). The direction and amount of regional dip of the strata in northern and central Illinois vary. In the vicinity of the project area, the dip is gently southwest on the flank of the Downs Anticline at about 25 to 30 feet per mile. 2.5.1.1.4.1.1 Illinois Basin The Illinois Basin is an oval-shaped basin with the major axis trending approximately N 25

° W. The major axis of the basin is approximately 350 miles long, and the minor axis is approximately 250 miles long (see Figures 2.5-8 and 2.5-9). The deepest part of the basin, the Fairfield Basin, is in southeastern Illinois. Sediments in the Fairfield Basin are 12,000 to 14,000 feet thick

(Reference 19). In a purely structural sense, the Illinois Basin could be said to extend out to the axes or crests of the bounding arches (Reference 20). Strata in the Illinois Basin rises gently north at an average of 1° or less to the Wisconsin Arch. To the northeast, the Illinois Basin is separated from the Michigan Basin by the Kankakee Arch. To the east, the Illinois Basin rises gently to the CPS/USAR CHAPTER 02 2.5-12 REV. 11, JANUARY 2005 Cincinnati Arch. To the south, the Illinois Basin rises to the Pascola Arch (which is outside of the regional area). To the southwest, the Illinois Basin is bordered by the Ste. Genevieve Fault Zone. To the west and northwest, the Illinois Basin is bordered by the Lincoln Anticline and the Mississippi River Arch. The site is located in the central portion of the Illinois Basin, north of the area of greatest structural depression. The Illinois Basin began to form in Cambrian time and continued to develop intermittently until the end of the Paleozoic (Reference 2). The center of thickest deposition of the basin migrated northward during the Paleozoic (Reference 13). 2.5.1.1.4.1.2 Wisconsin Arch and Kankakee Arch The Wisconsin Arch is a south-to southeast-trending extension of the Wisconsin Dome. It can be traced into Illinois to the vicinity of the city of Kankakee where it appears to connect with the Kankakee Arch of Illinois and Indiana (Reference 15). The Wisconsin Arch has a Precambrian core and is believed to be the result of crustal uplift, whereas the Kankakee Arch acquired its structural relief chiefly by greater subsidence of the structural basins which lie on either side of

it. 2.5.1.1.4.1.3 La Salle Anticlinal Belt The La Salle Anticlinal Belt is more than 200 miles long and extends from a point north of the Illinois River, near La Salle, to the Indiana State line on the Wabash River south of Vincennes.

Its closest approach is 15 to 20 miles to the east of the site. A portion of the La Salle Anticlinal Belt is reflected in the regional bedrock geology map as a finger of rocks of varied ages, from Silurian through Mississippian, flanked by younger Pennsylvanian units (see Figure 2.5-3). This feature is shown on the regional geologic sections (Figure 2.5-4). The La Salle Anticlinal Belt is a complex structure consisting in many places of en echelon north-south trending folds and troughs (Reference 18). Structures of the La Salle Anticlinal Belt are commonly asymmetrically folded to the west and are nearly monoclinal. Dips on the west flank of the belt may be up to 2000 feet per mile (approximately 20

°), while the eastern flank is generally much more gently dipping, 25 to 50 feet per mile (approximately 1/2

°) (Reference 21). Several en echelon folds and troughs of the La Salle Anticlinal Belt within 50 miles of the site have been named. To the north and east some 5 to 10 miles from the site, is a small flexure trending parallel to the La Salle Anticlinal Belt known as the Downs Anticline. Several domes are present along this structure and lie to the north and east of the site. These features are discussed further in Subsection 2.5.1.2.3. A series of subsidiary structures of the La Salle Anticlinal Belt lie 15 to 35 miles southeast of the station site. These structures, beginning with the most westerly, are the Mattoon and Tuscola Anticlines and the Murdock and Marshall Synclines (Figure 2.5-9). They are all minor structures associated with the La Salle Anticlinal Belt within the Illinois Basin and they all trend nearly parallel to the La Salle Anticlinal Belt. The Mattoon Anticline is a small positive structure. The Tuscola Anticline plunges southeastward and its west limb forms the west limb of the La Salle Anticlinal Belt. The Murdock Syncline plunges gently southward and is adjacent to the east limb of the Tuscola Anticline. The Marshall Syncline is asymmetrical, with a steep west flank, and plunges gently southward. The Ashton Arch (Figures 2.5-8 and 2.5-9) may be an extension of the La Salle Anticlinal Belt.

CPS/USAR CHAPTER 02 2.5-13 REV. 11, JANUARY 2005 Initial deformation along the La Salle Anticlinal Belt took place in Mississippian or pre-Mississippian time. Deformation probably continued intermittently until the close of Paleozoic time (Reference 22). 2.5.1.1.4.1.4 Mississippi River Arch The Mississippi River Arch is a broad, corrugated fold which parallels the Mississippi River, with contiguous parts in Illinois, Iowa, and Missouri. It is located approximately 120 miles west of the site (Figures 2.5-8 and 2.5-9). The western flank of the arch gently subsides into the Forest City Basin (Reference 17). Development of the arch began in Mississippian time and continued into the Pennsylvanian, as indicated by the thinning of sedimentary strata which rise onto the arch from adjoining basins (Reference 14). The arch was probably subjected to additional deformation at the close of the Paleozoic (Reference 22). 2.5.1.1.4.1.5 Lincoln Anticline The Lincoln Anticline trends approximately northwest-southeast for 165 miles through eastern Missouri and western Illinois (Figures 2.5-8 and 2.5-9), and separates the Forest City Basin (which lies outside of the regional area) from the Illinois Basin. The fold is not a simple anticlinal structure, but rather a regional uplift upon which are superimposed domes, anticlines, synclines, and faults. The structure is asymmetric and has a maximum structural relief of 1000 feet (Reference 17). The fold is bounded on the south by the Cap Au Gres Faulted Flexure. The southwest side of the fold is marked by comparatively steep dips and faulting, while the northeast flank is marked by gentle dips and the absence of faulting. The Lincoln Anticline is believed to have formed intermittently from Ordovician to post-Pennsylvanian time (References 16, 23). The main period of development of the fold was in the interval from Late Mississippian to Early Pennsylvanian time (Reference 23). 2.5.1.1.4.1.6 Ozark Uplift The central part of the Ozark Uplift or Dome is located approximately 200 miles southwest of the site. The Ozark Uplift is the major structural feature in Missouri, and is a broad, slightly asymmetrical, quaquaversal fold (Reference 17). The structural center of this uplift is in Iron County, Missouri. The topographic axis extends from Barry to Iron Counties. The boundaries of the Ozark Uplift are somewhat vague in areas, particularly to the north and northwest; however, they generally correspond to the Ordovician-Mississippian rock contacts to the east and west and to the Mississippi Embayment to the south. The Ozark Uplift is separated from the Illinois Basin by the Ste. Genevieve Fault Zone. The Ozark uplift underwent uplift intermittently in Paleozoic, Mesozoic, and Tertiary time and perhaps into the present as well (Reference 17). 2.5.1.1.4.1.7 DuQuoin Monocline The DuQuoin Monocline is located in southwestern Illinois (Figure 2.5-9) where it forms the western boundary of the Fairfield Basin. The monocline has a north-south strike, and dips to the east. The structure extends from the vicinity of DuQuoin to a point approximately 20 miles north of Centralia, a total distance of about 60 miles (Reference 24). The monocline developed during subsidence of the Illinois Basin. There is no evidence of post-Paleozoic movement.

CPS/USAR CHAPTER 02 2.5-14 REV. 11, JANUARY 2005 2.5.1.1.4.1.8 Salem and Louden Anticlines The Salem and Louden Anticlines (Figure 2.5-9) are north-south trending structural highs in the Fairfield Basin. The Salem Anticline extends from central Jefferson County to central Marion County in southern Illinois, and is approximately 20 miles in length. The Louden Anticline is located 7 miles northeast of the Salem Anticline. The Louden Anticline extends from the northern county line of Marion County through east-central Fayette County, Illinois, and is approximately 35 miles long. See Illinois Geological Survey Circular 519. Pennsylvanian units thin over the Salem and Louden Anticlines, indicating that the two anticlines were uplifted during Pennsylvanian time and later (Reference 22). 2.5.1.1.4.1.9 Clay City Anticline The Clay City Anticline is a prominent structure in the Fairfield Basin. It trends north-south and has an axial trace approximately 57 miles long. The anticline is a semicontinuous series of anticlinal uplifts separated by saddles (Referenc e 25). DuBois and Siever (Reference 25) noted that the amplitude of the anticline increases with depth and decreases in the overlying Pennsylvanian strata. They interpreted this to imply that the structure developed during pre-Pennsylvanian time; however, the presence of the fold in the Pennsylvanian strata indicates some folding occurred during Pennsylvanian and/or, post-Pennsylvanian time. 2.5.1.1.4.1.10 Dupo-Waterloo Anticline The Dupo-Waterloo Anticline trends approximately N 20

° W from Monroe County, Illinois, through St. Louis, Missouri (Figure 2.5-9). The northern end of the anticline terminates against the Cap au Gres Faulted Flexure (Reference 22). Total structural relief of the anticline is at least 500 feet near Waterloo, Illinois. The western flank of the anticline dips much more steeply, 30° or more, than the eastern flank, which dips only 2

° or 3o° (Reference 22). Major movements along the anticline probably occurred from Late Mississippian to pre-Pennsylvanian time, with renewed uplift in post-Pennsylvanian and pre-Pleistocene time (Reference 22). 2.5.1.1.4.1.11 Pittsfield-Hadley Anticline The Pittsfield-Hadley Anticline trends northwest-southeast and crosses Lewis County, Missouri, and Adams and Pike Counties, Illinois (Figure 2.5-9). Pennsylvanian strata on the flanks of the anticline dip less steeply than those of the underlying Mississippian, suggesting an episode of uplift in the interval from post-Mississippian to pre-Pennsylvanian time, but the major period of uplift is considered to have been post-Pennsylvanian (Reference 22). 2.5.1.1.4.1.12 Cap au Gres Faulted Flexure The Cap au Gres Faulted Flexure, located some 117 miles southwest of the site, is a structure that extends continuously from western Pike County, Missouri, southeastward toward Lincoln County, then east across southern Calhoun County, Illinois, and into southwestern Jersey County, where it disappears beneath the broad alluvial valley of the Mississippi River.

Throughout its length, the flexure is a narrow zone along which the rocks dip steeply southward or southwestward. The total uplift of "structural relief" along the flexure averages about 1000 feet but it varies from place to place (Reference 26). Deep faulting has been inferred on the basis of steep dips although the surface strata do not appear to be faulted. The principal folding

of the Cap au Gres Flexure was post-St. Louis (Mississippian) and pre-Pottsville CPS/USAR CHAPTER 02 2.5-15 REV. 11, JANUARY 2005 (Pennsylvanian) (Reference 26). Later periods of movement may have occurred; however, there is no evidence of deformation of nearby Pleistocene deposits. 2.5.1.1.4.1.13 Sangamon Arch The Sangamon Arch is located in central and western Illinois (Figure 2.5-9). The crest of the arch trends northeast-southwest across the broad shelf area west of the Illinois Basin and toward the northern center of the Illinois Basin. Buschbach (Reference 22) defines the arch by the zero isopach of the Cedar Valley Limestone (Middle Devonian). The Sangamon Arch was formed by uplift during the Devonian and Early Mississippian. The arch is a relict structure that has been masked by post-Mississippian movement (Reference 22). A recent study by Calvert (Reference 27) has questioned the existence of the Sangamon Arch, however the structure is still recognized by the Illinois State Geological Survey. 2.5.1.1.4.1.14 Structures Associated with the Plum River Fault Zone Four minor structural features, all associated with the Plum River Fault Zone in Illinois, are located approximately 150 miles northwest of the site (Figure 2.5-11). There are similar type structures in adjacent Iowa. Successively from west to east, the Illinois structures are the Uptons Cave Syncline, the Forreston and Brookville Domes, and the Leaf River Anticline. The Forreston and Brookville Domes were previously considered to be a single domal structure called the Brookville Dome until subsequent drilling indicated the presence of two domal structures. All four of these minor structures, and their conterparts in Iowa, are considered to be associated with the development of the Plum River Fault Zone (References 28 and 29). The Plum River Fault Zone formed in the interval from post-Silurian to pre-Pleistocene time. Based on regional geologic history, the Plum River Fault Zone probably developed at the same time that major movements were occurring on other structures in the region, which was near the beginning of Pennsylvanian time and again in post-Pennsylvanian time (Reference 28). In Illinois, Illinoian Strata of the Plum River Fault Zone shows no evidence of tectonic disturbance

(Reference 30); in Iowa there is no known evidence of displacement of Pleistocene deposits (See Iowa Geological Survey letter in Attachment D2.5). 2.5.1.1.4.1.15 Moorman Syncline The Moorman Syncline is located in southeastern Illinois and northwestern Kentucky, approximately 170 miles south-southeast of the site. It trends roughly east-west, parallel to and south of the Rough Creek Fault Zone. The Moorman Syncline is smaller and narrower than the Fairfield Basin, but somewhat deeper. The major movement on the Moorman Syncline took place in post-Pennsylvanian time (Reference 20). The Moorman Syncline is also known as the

Rough Creek Graben. 2.5.1.1.4.1.16 Folds in Wisconsin Within the regional area are two major folds in southwestern Wisconsin (Figure 2.5-9). The Meekers Grove Anticline trends east-west from Dubuque County, Iowa to Green County, Wisconsin. Its amplitude ranges from 100 to 200 feet. It is approximately 80+ miles long and varies in width from 3 to 5 miles. The north limb of the anticline dips much more steeply than the south limb.

CPS/USAR CHAPTER 02 2.5-16 REV. 11, JANUARY 2005 The Mineral Point Anticline is a complexly curved fold in southwestern Wisconsin. It is at least 70 miles long, ranges in width from 5 to 8 miles, and has an amplitude of between 100 and 170 feet. The Mineral Point Anticline is asymmetric with a more steeply dipping north limb. Both the Meekers Grove and Mineral Point Anticlines were probably formed in Late Paleozoic time (Reference 31). 2.5.1.1.4.1.17 Folds in Iowa Harris and Parker (Reference 32) have delineated five anticlines in southeastern Iowa by virtue of borehole data and structure contours on the Burlington Limestone (Mississippian). The five structures which generally parallel one another are the Bentonsport, Skunk River, Burlington, Sperry, and Oquawka Anticlines. Axial trends of these folds vary from N 55

° W to N 65

° W. Since 1964 the Iowa Geological Survey has reinterpreted the folded area (Attachment D2.5, Paragraph 3 of letter dated November 6, 1978) and concluded that the structures are not as continuous as shown. There is evidence for a series of northeast-trending folds cutting across the southeast-trending folds. The age of folding is difficult to establish with precision but it is thought to be Mississippian (Reference 32). 2.5.1.1.4.1.18 Minor Folds In Missouri Several minor folded structures are present in the southeastern Missouri portion of the regional area. These structures have formed at intermittent times during the Paleozoic. Reference 17 and the Clinton PSAR give more precise estimates. The Troy-Brussels Syncline has an east-west strike (Reference 17) and is located approximately 125 miles southwest of the site (Figure 2.5-9). It extends from near Troy, Missouri to west of Brussels, Illinois. Age of movement was Late Mississippian-Early Pennsylvanian. The Cuivre Anticline is located southwest of the Troy-Brussels Syncline (Figure 2.5-9). It has an axis which strikes N 80

° W, and it plunges to the southeast (Reference 17). It is separated from the Lincoln Anticline by the Troy-Brussels Syncline. Movement was post-Mississipian.

The Eureka-House Springs Anticline is located approximately 150 miles southwest of the site (Figure 2.5-9). The axis of the Eureka-House Springs Anticline trends approximately southeast-northwest, and appears to plunge both to the southeast, in Jefferson County, Missouri, and to the northwest, in St. Louis County, Missouri (Reference 17). Movement was post-Mississippian. The Farmington Anticline is located approximately 175 miles southwest of the site in Ste. Genevieve and St. Francois Counties, Missouri. The axis of the anticline trends N 30

° W, and the anticline extends for a distance of 15 to 20 miles. The northeast limb of the anticline is steeper with dips up to 4

°, while the dip on the southwest limb is 1

°. The anticline is terminated by faults at both ends. It has been concluded that the Farmington Anticline is no older than Devonian (Reference 17). The Plattin Creek Anticline is located approximately 160 miles southwest of the site. It trends approximately N 15

° E, and is 18 miles long (Reference 17). Movement of the structure is thought to be late Mississippian-Early Pennsylvanian.

CPS/USAR CHAPTER 02 2.5-17 REV. 11, JANUARY 2005 The Crystal City Anticline is located approximately 155 miles southwest of the site. The south limb dips at 5

° while the north limb has a more gentle 2

° dip (Reference 17). Movement is post-Mississippian. The Krugers Ford Anticline is located approximately 175 miles southwest of the site in Osage and Gasconade Counties, Missouri (Figure 2.5-9). The anticline strikes northeast-southwest with the steep flank on the southeast (Reference 17). Structural movement was post-

Ordovician. The Pershing-Bay-Gerald Anticline is located approximately 180 miles southwest of the site, and trends approximately northwest-southeast (Figure 2.5-9). It passes through the villages of Pershing and Bay, and near Gerald, Missouri to the faulted area near Anaconda (Reference 17). Structural activity was in post-Mississippian and early Pennsylvanian time. The Mineola Structure is located approximately 165 miles southwest of the site in Montgomery County, Missouri (Figure 2.5-9). The Mineola Structure is an asymmetric dome, trending northwest-southeast, with steeper dips on the southwest side (Reference 17). The structure was active in Pennsylvanian and post-Pennsylvanian time. The Auxvasse Creek Anticline is located approximately 175 miles southwest of the site (Figure 2.5-9). The Auxvasse Creek Anticline strikes northwest, and is asymmetric with a steep southwest limb (Reference 17). The anticline developed in post-Pennsylvanian time. The Brown's Station Anticline is located approximately 200 miles southwest of the site (Figure 2.5-9). It trends northwest-southeast. The southwest limb of the anticline is steep and may be faulted (Reference 17). The structure developed in late Mississippian or Pennsylvanian time. The Davis Creek Anticline is located in Audrain County, Missouri, approximately 180 miles from the site (Figure 2.5-9). The structure trends approximately northwest-southeast and was mapped using subsurface data (Reference 17). The structure was active in post-Mississippian

time. The Mexico Anticline is located in Audrain County, Missouri, approximately 175 miles from the site. This anticline strikes northeast-southwest approximately normal to the general northwest-southeast trending structures of the area (Figure 2.5-9). The anticline is late Pennsylvanian or

Pennsylvanian in age. The College Mound-Bucklin Anticline is located 200 miles west-southwest of the site (Figure 2.5-9). The anticline strikes northwest-southeast and has a gentle plunge to the northwest (Reference 17). It follows the pattern of other such flexures in northern Missouri in having a gentle northeast limb and a slightly steeper southwest limb. Structural development occurred in

late Pennsylvanian or post-Pennsylvanian time. 2.5.1.1.4.2 Faulting Major faults within a 200-mile radius of the site are discussed in the following subsections. Figures 2.5-8 and 2.5-10 show the location and extent of these faults. Table 2.5-2 summarizes the type and displacement of each fault, and gives the age of the last movement.

CPS/USAR CHAPTER 02 2.5-18 REV. 11, JANUARY 2005 2.5.1.1.4.2.1 Sandwich Fault Zone The Sandwich Fault Zone is located in northern Illinois, and strikes west-northwest from western Will County, Illinois to Ogle County, Illinois (References 22 and 30). The closest approach to the site is approximately 90 miles north-northeast (Figure 2.5-10). The northeast side of the fault zone is down thrown, to a maximum displacement of 800 feet (Reference 30). The fault zone forms the northern boundary of the Ashton Arch. Movements along the fault zone occurred in the interval between post-Silurian and pre-Pleistocene time. No rocks of intervening ages are present, which prevents better definition of the movements. However, major movements along the fault zone may have been contemporaneous with folding of the La Salle Anticlinal Belt during post-Mississippian to pre-Pennsylvanian time (References 22 and 30). There is no relationship between historical earthquake epicenters and the Sandwich Fault Zone

(Reference 30). 2.5.1.1.4.2.2 Plum River Fault Zone The Plum River Fault Zone, (formerly called the Savanna Fault and the Savanna Anticline) is a generally east-west trending zone of high angle faults extending from near Leaf River (Ogle County), Illinois, west-southwest into Southern Linn County, Iowa (References 28 and 29, Figures 2.5-10 and 2.5-11). The width of the fault zone is less than one-half mile, and vertical displacement along the fault zone is 100 to 400 feet, north side down (Reference 28). The age of movement in Illinois has been limited to post-middle Silurian to pre-middle Illinoian (Pleistocene) (Reference 28). However, there are no deposits representing the interval from middle Silurian to middle Illinoian (Pleistocene) time, and thus, more precise dating of the faulting is not possible. It seems likely that the Plum River Fault Zone formed at the same time as the La Salle Anticlinal Belt and Mississippi River Arch, near the beginning of Pennsylvanian time and again in post-Pennsylvanian time (Reference 28). In Iowa there is evidence of Paleozoic movement of the Plum River Fault Zone (Reference 29) but no known evidence of displacement of Pleistocene strata (Attachment D2.5, Paragraph 1 of letter dated November 6,

1978). Minor folded structures are associated with the Plum River Fault Zone. In Illinois, named folds are the Forreston Dome, the Brookville Dome, the Leaf River Anticline, and Uptons Cave Syncline. These structures are discussed in Subsection 2.5.1.1.4.1.14. 2.5.1.1.4.2.3 Cap au Gres Faulted Flexure The Cap au Gres Faulted Flexure is discussed in Subsection 2.5.1.1.4.1.12.

2.5.1.1.4.2.4 Centralia Fault The Centralia Fault is located in Marion and Jefferson Counties, Illinois. It is a zone of several parallel faults, striking north-south, which are parallel to, and 1 mile east of the DuQuoin Monocline (Reference 22). The closest approach of the fault is approximately 110 miles south of the site (Figure 2.5-10). Downthrow on the fault is 160 to 200 feet on the west (Reference 33). Although faults have been observed in several coal mines in the Centralia area, they have no surface expression (References 22 and 33). The faulting occurred in the interval from post-Pennsylvanian to pre-Pleistocene time (Reference 22).

CPS/USAR CHAPTER 02 2.5-19 REV. 11, JANUARY 2005 2.5.1.1.4.2.5 Ste. Genevieve Fault Zone The Ste. Genevieve Fault Zone trends approximately northwest-southeast from Perry County, Missouri to Union County, Illinois, and is located approximately 160 miles southwest of the site (Figure 2.5-10). The faults are high angle and form numerous horsts and grabens. The net displacement is down to the north and east with the maximum displacement greater than 1000 feet, and possibly up to 2000 feet. The Ste. Genevieve Fault Zone forms a sharp boundary, a few miles wide, between the Illinois Basin and the Ozark Uplift. Movement on the Ste.

Genevieve Fault Zone may have started as early as the Devonian, and movement is known to have occurred both in the post-Mississippian, pre-Pennsylvanian interval and in post-Pennsylvanian time (Reference 22). 2.5.1.1.4.2.6 Rough Creek Fault Zone Buschbach (Reference 22) states that the Rough Creek Lineament is a series of faults and fault zones extending generally east-west through western Kentucky and southern Illinois. In Kentucky, it includes the Rough Creek Fault Zone. In Illinois, it includes the east-west portion of the Shawneetown Fault Zone to the east, and the Cottage Grove Fault System to the west.

Heyl (Reference 34) suggests that strike-slip faulting or wrench faulting is a major component in the Rough Creek Lineament, however, this has been disputed. Heyl tentatively includes it in a line or zone of faults, monoclines, and igneous intrusions. The line extends east-west for 800 miles along the 38th parallel from West Virginia to at least as far west as the Ozark Uplift. In Illinois, the lineament includes numerous high angle reverse faults with the south side upthrown. They appear to be the result of compressional forces from the south, and they display evidence of some horizontal movements. The eastern part of the lineament, the Shawneetown Fault Zone, is dominated by high angle thrust faulting. Displacement is locally as great as 3400 feet and may be considerably more. The Shawneetown Fault Zone extends westward along the prominent hills in southern Gallatin County, curves southward around Cave Hill in Saline County, leaves the Rough Creek Lineament and joins the southwest trending Herod Fault to form the Lusk Creek Fault Zone. The western portion of the lineament, the Cottage Grove Fault System, appears to have formed at roughly the same time as the Shawneetown, but displacements are much diminished, with maximum displacements of about 250 feet. From all available evidence it appears that the age of faulting along the Rough Creek Lineament is chiefly post-Pennsylvanian, pre-Late Cretaceous, although some workers have suggested the possibility of later movements because of recent seismic activity in the general

area. 2.5.1.1.4.2.7 Wabash Valley Fault System The Wabash Valley Fault System is a series of generally parallel faults, 125 miles from the site at its closest approach, that is terminated at the Rough Creek Fault Zone (Reference 35). It extends north-northeastward from the Rough Creek Fault Zone generally parallel to the Wabash River in southeastern Illinois and southwestern Indiana (Figure 2.5-10). The faults are high angle, normal faults (Reference 22), some of which border horsts and grabens (Reference 36).

Maximum displacement known on the faults is a little over 400 feet, but displacements from a few to 200 feet are more common. Displacements of Mississippi and Pennsylvanian age strata along the faults is the same, indicating that the faults are post-Pennsylvanian in age. No displacement has been recognized in Pleistocene deposits, thus the faulting was concluded in pre-Pleistocene time (Reference 22).

CPS/USAR CHAPTER 02 2.5-20 REV. 11, JANUARY 2005 2.5.1.1.4.2.8 Northeast Trending Faults South of the Rough Creek Fault Zone South of the Rough Creek Fault Zone, in southeastern Illinois and western Kentucky, is an intensely faulted area (Figure 2.5-10). Faults in this area trend dominantly northeast-southwest and east-west. Both normal and reverse faults are found in this area (Reference 37), but most of the faults are normal faults (Reference 38). Some graben structures are known in the area (Reference 37). Displacements on faults in the area are variable (Reference 38), but displacements up to 2000 feet have been reported (Reference 37). Faults south of the Rough Creek Fault Zone are generally post-Pennsylvanian in age and terminate beneath the Cretaceous cover of the Mississippi Embayment, but there is some evidence that faulting on a small scale continued into the Tertiary (Reference 14). 2.5.1.1.4.2.9 Mt. Carmel Fault The Mt. Carmel Fault trends approximately N 25

° W across south central Indiana from Washington County, north to Monroe County, approximately 130 miles southeast of the site (Figure 2.5-10). It is a normal fault, downthrown to the west, with a displacement which may be in excess of 200 feet. Movement along the fault may have begun in late Mississippian time and probably was concluded by Early Pennsylvanian time (Reference 39). 2.5.1.1.4.2.10 Fortville Fault The Fortville Fault trends northeast-southwest through central Indiana, approximately 145 miles east of the site (Figure 2.5-10). The Fortville Fault is a normal fault and is about 54 miles long. On the basis of structure contours on the top of the Ordovician Trenton Limestone (Reference 40), the fault has a vertical displacement of about 60 feet, downthrown to the southeast (Reference 41). The fault displaces Devonian age strata, but does not displace overlying Pleistocene age deposits (Reference 42). 2.5.1.1.4.2.11 Royal Center Fault The Royal Center Fault trends northeast-southwest for 47 miles in north-central Indiana. It is approximately 125 miles from the site at its closest approach (Figure 2.5-10). On the basis of structure contours on the top of the Ordovician Trenton Limestone (Reference 40), the Royal Center Fault is a normal fault, downthrown to the southeast, with a vertical displacement of about 100 feet (Reference 41). The fault displaces Devonian age strata, but does not displace overlying Pleistocene age deposits (Reference 42). 2.5.1.1.4.2.12 Localized Faults in Wisconsin Locally restricted faults are known in the regional area from two locales in southern Wisconsin: southern Dane County, Wisconsin, and at Waukesha, Wisconsin. These faults are not shown in Figure 2.5-10. The locally restricted faults reported in southern Dane County, Wisconsin trend northeast. Displacements on these faults in the southern Madison, Wisconsin area are approximately 150 feet, and in the Mt. Vernon, Wisconsin area, about 60 feet (Reference 43). Displacements in the general southern Dane County area may be as great as 300 feet with the northwest side down (Reference 44). The age of the faulting in southern Dane County is post-Ordovician pre-Pleistocene (References 43, 44). These faults in southern Dane County do not correlate with CPS/USAR CHAPTER 02 2.5-21 REV. 11, JANUARY 2005 the regionally continuous Madison fault reported by Thwaites (Subsection 2.5.1.1.4.2.15), the existence of which is now considered doubtful by present workers (Reference 43). Faulting is known at an exposure in Waukesha, Wisconsin, where the southeast side of the fault has been downthrown 27 feet. Movement on the fault is considered to have occurred in post-Silurian pre-Pleistocene time. Subsurface data does not substantiate extending this fault past the Waukesha town limits. This fault was originally the basis of an inferred fault in the Precambrian basement (the "Waukesha" Fault, described in Subsection 2.5.1.1.4.2.15), but recent studies reject the interpretation of a major fault extending from the Waukesha area (Reference 43). 2.5.1.1.4.2.13 Cryptovolcanic or Astrobleme Structures Cryptovolcanic or astrobleme structures, shown in Figures 2.5-8 and 2.5-10, include the Glasford, Kentland, and Des Plaines Disturbances. These are complex, local areas of intense shattering, thought to be due to meteorite impact. The Glasford Disturbance is approximately 60 miles northwest of the site. The Glasford Disturbance has been outlined by gravity surveying and structural drilling; the disturbance consists of a dome with a normal sequence of Upper and Middle Paleozoic strata underlain by a series of jumbled blocks in a breccia matrix of Cambrian formations uplifted about 1000 feet.

Middle and Lower Ordovician strata are not recognizable in the disturbed zone. The chaotic condition of the pre-Upper Ordovician rocks suggests that this disturbance was caused by a violent explosion which probably took place in early-Late Ordovician time (Reference 45). This feature is 2-1/2 miles in diameter and is very local in nature. The Glasford Disturbance is the closest structure of this type to the site. The Des Plaines Disturbance is located in northeastern Illinois, approximately 140 miles from the site (Figure 2.5-10). The disturbance is roughly circular, about 5-1/2 miles in diameter, and contains many small blocks separated by normal and reverse faults. A graben, in which Mississippian and Pennsylvanian age rocks are preserved, partly surrounds a central uplifted core. The Des Plaines Disturbance was formed sometime in the interval from post-Pennsylvanian to pre-Wisconsinan (Pleistocene) time (Reference 46). The Kentland Disturbance is located in western Indiana, approximately 90 miles from the site. As in other cryptovolcanic structures, the Kentland Disturbance is characterized by a nearly circular outline, a central uplift, and a marginal, ring-shaped depression with irregular and local faulting (Reference 15). The Kentland Disturbance is considered to be late Paleozoic or Mesozoic in age. 2.5.1.1.4.2.14 Postulated Faults in Illinois 2.5.1.1.4.2.14.1 Oglesby Fault and Tuscola Fault From maps of the Trenton (Ordovician) structure in Ohio, Indiana, and northern Illinois, Green (Reference 47) postulated two faults in Illinois, the Oglesby Fault and the Tuscola Fault (Figure 2.5-10). These two faults approximately coincide with the western flank of the La Salle Anticlinal Belt, and the western sides of these faults were inferred to be downthrown. The Illinois State Geological Survey has never accepted the existence of the Oglesby and Tuscola Faults, and has found no evidence to support such an interpretation. The Illinois State Geological Survey contends that the differences in elevations of the top of the Trenton are not CPS/USAR CHAPTER 02 2.5-22 REV. 11, JANUARY 2005 due to faulting, as suggested by Green, but are simply due to dipping beds with dips of a few to 10° (Reference 48). 2.5.1.1.4.2.14.2 Chicago Area Basement Fault On the basis of gravity and seismic geophysical evidence, McGinnis postulated a basement fault zone in the metropolitan Chicago area north of and about paralell to the Sandwich Fault Zone (Figure 2.5-10). It was inferred that the southwest side of this fault zone was downthrown up to 900 feet. The presence of this basement fault has not been verified. McGinnis (Reference 49) has postulated that movement on this basement fault was completed by Middle Ordovician time. 2.5.1.1.4.2.14.3 Chicago Area Minor Faults As a result of a recent seismic survey in the metropolitan Chicago area, 25 faults were inferred, with displacements up to 55 feet. None of these faults involves wide shear zones or detectable scarps on the rock surface (Reference 50).

Other faults that have been observed in natural outcrops and quarries in the Chicago area have displacements from a few inches to a few feet, but most show less than 1 foot of movement (Reference 50). 2.5.1.1.4.2.15 Postulated Faults in Wisconsin Thwaites' map of the buried Precambrian surface in Wisconsin (Reference 51) postulates the existence of several faulted areas in southern Wisconsin. Ostrom (Reference 43) stated that Thwaites' map is diagrammatic and does not represent a detailed study of each fault. Three of the postulated faults, named by Dames & Moore as the Janesville, Madison, and Waukesha Faults, have a noticeable difference in the elevation of the Precambrian basement, which was interpreted by Thwaites to be the result of faulting. It is now believed (Reference 43) that this difference in the elevation of the Precambrian basement is not due to faulting, but due to topographic relief on an erosional Precambrian basement surface. 2.5.1.1.5 Gravity and Magnetic Anomalies Gravity anomalies are caused by lateral rock density changes which may result from: a. structure, unconformities, and lithologic changes in the sedimentary rocks;

b. igneous intrusives;

c. relief on the crystalline basement surface; and

d. lithologic changes in the crystalline portion of the earth's crust and upper mantle. Figure 2.5-12 represents the Bouguer gravity anomaly map of the region surrounding the site. The anomaly pattern north of 38.5-39

° is typically midcontinental in style (Reference 52). The site lies on the flank of a moderate gravity low associated with the La Salle Anticlinal Belt where dense Precambrian crystalline rocks lie at relatively shallow depths beneath the surface (Reference 2).

CPS/USAR CHAPTER 02 2.5-23 REV. 11, JANUARY 2005 Magnetic anomalies are largely caused by lateral changes in the concentration of the mineral magnetite. Therefore, in the site region, magnetic anomalies result from: a. igneous intrusives; b. relief on the crystalline basement surface; and c. lithologic changes in the crystalline portion of the earth's crust above the Curie point isotherm. Detailed aeromagnetic surveys have been carried out in Illinois only south of 39

° N (References 54 and 55). An early ground magnetic survey (Reference 56) is shown in Figure 2.5-13. The scale of the gravity anomaly map does not permit anything other than a general correlation between gravity anomalies and major regional structures. 2.5.1.1.6 Regional Groundwater Groundwater conditions in the vicinity of the site are discussed in Subsection 2.4.13.

2.5.1.1.7 Man's Activities A discussion of man's activities in the site area is included in Subsection 2.5.1.2.7. 2.5.1.2 Site Geology 2.5.1.2.1 Site Physiography The site lies within the Bloomington Ridged Plain physiographic subsection of the Till Plains Section (Figure 2.5-1). The main plant is located in an area of uplands, consisting of Wisconsinan-age ground moraine, that have been dissected by the Salt Creek and the North Fork of the Salt Creek (Figures 2.5-14 and 2.5-15). The ultimate heat sink and associated structures are located along the North Fork of the Salt Creek (Figures 2.5-14 and 2.5-16), and the main dam is located below the confluence of the Salt Creek and the North Fork of the Salt

Creek (Figure 2.5-14). The uplands consist of gently rolling ground moraine, located just east of the Shelbyville end moraine, with local relief of about 10 feet, except near the drainageways. Average elevation of the uplands is approximately 740 feet MSL. Two perennial streams, Salt Creek and North Fork of the Salt Creek are present in the site area. The two streams join in the southern portion of the site area (Figure 2.5-14). The two streams flow generally to the southwest with gradients of 2 to 3 feet per mile in the site area. They have eroded through the upland deposits of the Wisconsinan-age Wedron Formation and Robein Silt, the Illinoian-age weathered Glasford Formation, and into the upper part of the Illinoian-age unaltered Glasford Formation. The elevation of the floodplains of the two streams in the site area is at approximately 660 feet MSL. Maximum relief in the site area is on the order of 80

feet.

CPS/USAR CHAPTER 02 2.5-24 REV. 11, JANUARY 2005 2.5.1.2.2 Site Stratigraphy The general stratigraphy of the Clinton Power Station (CPS) site is shown in Figure 2.5-18, Site Stratigraphic Column and a detailed description of stratigraphic units exposed in plant excavation is presented in Attachment C2.5. Boring logs are presented in Figures 2.5-19 through 2.5-270. Topography, plant structures, plot plans, and location of geologic sections across the site are given in Figures 2.5-271, 2.5-272, 2.5-273, and 2.5-16. The strata underlying the site consist of overburden deposits, about 225 to 360 feet in thickness in the upland areas, resting on Pennsylvanian-age bedrock. The overburden materials, in order of increasing age, consist of stream alluvium, windblown loess, and glacial drift. Colluvium and glacial outwash are also present. The stratigraphic nomenclature of the overburden deposits used for the excavation mapping reports (Attachment C2.5) and the FSAR is different from that used for the PSAR. This change was brought about in order that the nomenclature of the site glacial drift would be consistent with that used by the Illinois State Geological Survey. Figure 2.5-274 is a chart showing the correlation between

stratigraphic terms used in the PSAR and those used in the FSAR and excavation mapping reports. The terminology changes do not indicate any difference between the lithologic units encountered in the borings and those encountered in the excavations. Overburden materials and bedrock deposits are discussed in greater detail in the following subsections. 2.5.1.2.2.1 Overburden Materials Overburden materials occurring in the site vicinity, in order of increasing age, consist of stream alluvium, windblown loess, and glacial drift. Colluviam and glacial outwash are also present. Figures 2.5-275 through 2.5-280 show cross sections which portray the relationships of the various overburden materials across the site area, as determined from the boring logs.

Attachment C2.5 describes the overburden materials exposed in excavations for the Clinton

Power Station. Addressing valley deposits the stream alluvium and recent channel deposits, known as the Cahokia Alluvium, are composed of poorly sorted silt, clay, and silty sand, with sand and gravel lenses. Beneath the alluvium are glacial outwash deposits of the Henry Formation consisting of yellow-brown fine to coarse sand and gravel, pockets of silty-clayey material, and a basal lag gravel. Not differentiated in field mapping, these deposits range up to 35 feet in thickness, collectively, in the site borings and excavations, and are restricted to the valleys of Salt Creek and North Fork of Salt Creek. Also in the valleys, at the base of the valley walls, are local deposits of Peyton Colluvium directly overlying the Cahokia Alluvium and, in-places, Illinoian Till of the Glasford Formation. The colluvium is composed of brown clayey silts with minor amounts of gravel. In the excavations for structures for the Clinton Power Station, (in the valleys), the alluvium-outwash colluvium was underlain by unaltered Illinoian-age Glacial Till of the Glasford Formation. The Cahokia Alluvium, Henry Formation, and Peyton Colluvium were referred to as the Salt Creek Alluvium, Flood Plain Alluvium, or Recent Channel Deposits in the PSAR text and as the Salt Creek Alluvium on the boring logs (Figure 2.5-274). The Cahokia Alluvium is Holocene and possibly, in part, Valderan/Twocreekan in age; the Henry Formation is Woodfordian (probably early) in age.

CPS/USAR CHAPTER 02 2.5-25 REV. 11, JANUARY 2005 In the uplands between drainages, loess and glacial drift comprise the surficial deposits. Total drift thickness in the uplands varies from 210 to 310 feet in borings, this difference being largely due to the relief on the bedrock surface beneath the glacial drift (Figures 2.5-7 and 2.5-281). The loess, known as the Richland Loess, consists, generally, of a brown clayey silt with a trace of sand, and is present in the uplands at the site to thicknesses of 5 to 10 feet. A modern soil profile has been developed in the Richland Loess. Both the Richland Loess and the Wedron Formation which it overlies were deposited during the Wisconsinan Stage of glaciation. The Richland Loess was referred to as "Loess" in the PSAR text and on the boring logs. The Richland Loess may be Holocene in age, in part. The deposits of glacial drift in the site area form a complex sequence of materials (Figure 2.5-18). The uppermost deposits confined to the upland consist of Wisconsinan-age glacial till of the Wedron Formation. The Wedron Formation was referred to as Wisconsinan Glacial Till in the PSAR and boring logs (Figure 2.5-274). The Wedron Formation is from 20 to 55 feet thick in the site area where it has not been partially removed by erosion. It is composed of stiff to very stiff clayey sandy silt till, that is brown in cloration in the upper oxidized zone but grades to gray in the unoxidized zone. Discontinuous lenses of stratified sand, silt or gravel are randomly interbedded within the till of the Wedron Formation at the CPS site. Underlying the till of the Wedron Formation is the Robein Silt. The Robein Silt, also restricted to the uplands, was deposited during the Farmdalian Substage of the Wisconsinan Stage. It is a dark colored silt, rich in organic material. It is present over much of the site area, and may be up to 2 feet thick, although locally it may be absent due to erosion. The Farmdale Soil (Reference 5) is developed in the Robein Silt. The Robein Silt was previously included with the "interglacial zone" in the boring logs and the "Sangamonian interglacial soil zone" or "Sangamon soil interval" in the PSAR text (Figure 2.5-274). Underlying the Robein Silt, and also present under the valleys, are deposits of Illinoian-age collectively referred to as the Glasford Formation. The upper part of the Glasford Formation was weathered during the late Illinoian Sangamonian and possibly Altonian stages, and these weathered deposits are referred to as the weathered Glasford Formation in the USAR and in the excavation mapping reports (Attachment C2.5) and as Interglacial Zone, Sangamon Interglacial Zone, or Sangamon Soil Interval in the PSAR. Preserved mostly in the uplands, the weathered Glasford Formation is leached, characteristically black, dark brown, green or bluish-green and is 10 to 15 feet thick in the site area. The weathered materials are dominantly glacial till, consisting of silty clay and clayey silt but locally they may be discontinuous lenses of silts, sands, or sandy silts interbedded within the glacial till of the Glasford Formation. The boundary between the weathered Glasford Formation and the unaltered Glasford Formation is marked by the occurrence of calcareous glacial till of the Glasford Formation. The unaltered Glasford Formation at the site ranges in thickness from 90 to more than 140 feet. It is

dominantly a hard, gray-brown sandy silt till. Discontinuous layers of stratified sand, gravel, or silt, up to 2 to 3 feet in thickness may be interbedded within the till in the uppermost part of the unaltered Glasford Formation at the site (Attachment C2.5). The lower part of the unaltered Glasford Formation exposed in the excavations appears to have virtually no interbedded stratified deposits. Excavations for plant structures extended down into only the unaltered Glasford Formation, and data on the underlying deposits are from borehole samples. The unaltered Glasford Formation was referred to as "Illinoian Till" or "Illinoian Glacial Till" in the PSAR text and on the boring logs (Figure 2.5-274).

CPS/USAR CHAPTER 02 2.5-26 REV. 11, JANUARY 2005 Beneath the Glasford Formation is a complex assemblage of glacial materials consisting of gray to brown clay till (which is occasionally sandy), reworked till and outwash, and glacio-lacustrine gray silt. Correlation of these formations throughout the site area is difficult and uncertain. The sequence is probably pre-Illinoian in age and varies in thickness from 10 to 105 feet. These materials have been tentatively assigned to the Banner Formation which was deposited during the Kansan Stage (Figure 2.5-18). (Existing classifications of pre-lllinoian Plastocene glacial deposits have been questioned by some as noted in Reference 4). In some areas of the site, as beneath the main power block, the complex of probable pre-Illinoian till, outwash, and glaciolacustrine deposits lies in direct contact with bedrock.

Generally, however, it is underlain by a clean sand and gravel deposit of highly variable thickness which is identified as Kansan Stage glacial outwash (Mahomet Sand Member of the Banner Formation). This interval shows pronounced thickening where the bedrock surface slopes to lower elevations and is a glaciofluvial filling in the bedrock valleys. Its thickness ranges from zero on the highest bedrock surfaces to 140 feet at the lowest bedrock elevations. 2.5.1.2.2.2 Bedrock Formations Bedrock was penetrated by 19 of the borings. Twelve borings were located in the area of the power block (Figure 2.5-271), two at or near the dam site (Figure 2.5-272), and three in the ultimate heat sink area (Figure 2.5-16). One boring was drilled about 2 miles northeast of the station site for the purpose of identifying the regional trend of drift thicknesses and bedrock lithologies. The site is underlain by bedrock of Pennsylvanian age. The bedrock surface at the site is an erosional surface that varies in elevation from 360 to 510 feet MSL. Figure 2.5-281 shows the general configuration of the bedrock surface in the Dewitt-McLean County area as determined from boring logs and geophysical data. Data from borings for the CPS site have necessitated some modifications to the location of the bedrock valleys that are shown in the site area in Figure 2.5-281. The bedrock surface in the site area, as determined from these borings, is shown in Figures 2.5-17 and 2.5-282. The Pennsylvanian bedrock beneath the CPS site is characterized by sharp vertical changes in rock type and by laterial persistence of units such as limestones or coals, where they have not been removed by erosion. Regional marker beds which were encountered during the site investigation are the Shoal Creek Limestone Member, the No. 8 Coal Member, and the No. 7 Coal Member; these units were identified by the Illinois State Geological Survey. The uppermost Pennsylvanian strata in the site area belong to the Bond Formation of the McLeansboro Group. The Shoal Creek Limestone Member is a marker bed at the base of the Bond Formation. The Shoal Creek Limestone Member is found beneath the power block and ultimate heat sink in areas where the Pennsylvanian bedrock has not been eroded below an elevation of 495 feet MSL. The Shoal Creek Limestone Member is a fine to coarse crystalline limestone with irregular shale partings; in the upper portion of the unit, there are numerous open and clay-filled weathered bedding planes. Underlying the Bond Formation is the Modesto Formation which is also part of the McLeansboro Group. Three distinctive units of the Modesto Formation were found in the site area. The upper part of the Formation contains an unnamed limestone unit that is continuous across the CPS site in areas where the bedrock has not been eroded below an elevation of 472 feet MSL. This limestone is an argillaceous, fine crystalline limestone, which is variable in thickness, and CPS/USAR CHAPTER 02 2.5-27 REV. 11, JANUARY 2005 contains interbedded shale. The No. 8 Coal Member of the Modesto Formation is 1 foot thick at the CPS site, and was encountered at elevation 431 feet MSL in Boring P-38. Below elevation 424 feet MSL at the site, the Modesto Formation is predominantly siltstone and shale. A limestone bed was encountered in Boring D-11, however, at an elevation of 360 feet MSL. Underlying the McLeansboro Group is the Kewanee Group, which is also of Pennsylvanian age.

The uppermost formation in the Kewanee Group is the Carbondale Formation whose top is marked by the No. 7 Coal Member. This unit is 2.5 to 3 feet thick in the site area and was encountered and correlated in the three borings across the site. A chemical analysis was performed on samples from Boring P-38 of the No. 7 and No. 8 Coal Members by the Illinois State Geological Survey, and the results are presented in Table 2.5-3. Information on Pennsylvanian strata below the No. 7 Coal Member and on older pre-Pennsylvanian rocks was derived from a deep boring (1,621 feet) drilled approximately one mile west of the site (NW part of Section 27, T.20N., R.3E). About 440 feet of Pennsylvanian-aged strata, consisting principally of cyclothems, lie below the No. 7 Coal Member. Below the Pennsylvanian, the boring penetrated some 560 feet of Mississippian shale and limestone, which were in turn underlain by 180 feet of Devonian limestone and shale. The drill hole terminated in Silurian dolomite. Regional correlation between this boring, other deep borings in the vicinity, and borings at the site is shown in Figure 2.5-425. See Subsection 2.5.1.2.5 for information on older strata beneath the site. 2.5.1.2.3 Site Structural Geology The site is located to the southwest of the Downs Anticline, a subsidiary fold to the LaSalle Anticlinal Belt (Figure 2.5-285). The Downs Anticline trends south to south-southeastward from a point approximately 10 miles north of Bloomington, Illinois to a point approximately 4 miles south of DeLand, Illinois. At its closest the anticline is 4 miles northeast of the site (Figure 2.5-285). Several small domes are located along the axis of the Downs Anticline (Reference 57).

Three of these domes lie between 5 and 10 miles from the site: the Wapella East Dome to the northwest, the Parnell dome to the northeast, and the DeLand Dome to the southeast. The Downs Anticline is an asymmetrical, almost monoclinal fold. Cross sections through the Downs Anticline are shown in Figures 2.5-286 and 2.5-283. Although these figures have considerable vertical exaggeration in order to show the structure, the tops of Trenton (Ordovician) and Hunton (Silurian-Devonian) dip only at about 3

° and 2°, respectively, and the dips on the No. 2 Coal Member and the No. 7 Coal Member are even less (about 1/2

° to 1°). Rocks below and including the Hunton (Silurian-Devonian) were folded between post-Middle Devonian and pre-Pennsylvanian time, as shown by the accentuated bending of tops of the Trenton (Ordovician) and Hunton (Silurian-Devonian) in Figure 2.5-283. Rocks below and including the coal beds were folded between Pennsylvanian and pre-Pleistocene time. The Pleistocene stratum (Robein Silt) shown in Figure 2.5-283 does not reflect the folding of the coal beds. No tectonic folding or faulting was observed in the Pleistocene deposits exposed in the excavations at the CPS site, including the Robein Silt. The irregularity shown on the Pleistocene stratum shown in Figure 2.5-283 is due to deposition on an uneven erosional

surface. The bedrock surface is an erosional surface, and in the site area there is no general relationship between Paleozoic structures and bedrock topography. However, the Downs Anticline is located over a broad bedrock high, and a small circular bedrock high coincides with the Wapella East Dome. The relation of the other structural domes to bedrock topography is not as obvious.

CPS/USAR CHAPTER 02 2.5-28 REV. 11, JANUARY 2005 The DeLand Dome is located in broad bedrock uplands and the Parnell Dome is located on the flank of a bedrock valley. The location of these structures with respect to bedrock topography is shown in Figure 2.5-281. In some cases, structures and bedrock topography coincide; whereas in others, there is no relationship. Structure cannot, therefore, be inferred from bedrock topography. Borings at the site encountered small elevation differences in certain lithologic units of the bedrock as shown in Figures 2.5-275, 2.5-276, 2.5-279, and 2.5-284. Measured dips on these stratigraphic horizons corresponding to such elevation differences are less than 1

° (Figure 2.5-283) and average about 1/2

°. In the heat sink area, a dip of 1.9

° to the south was determined. A vertical exaggeration of 16 to 1 on the cross sections gives an exaggerated impression as to the magnitude of dip. The bedding slopes away from a high point beneath the site, and this possibly could be indicative of a minor structure. It is more probable, however, that this condition reflects normal irregularities typical of sedimentary contacts, even in tectonically undisturbed regions. Lensing and lithologic interfingering could easily account for such variations without the mechanism of uplift, compressional forces, or faulting. Borings D-31, H-6, and P-38 were drilled into or through the No. 7 Coal Member. The elevation of the coal in each boring correlates reasonably well with the structure contour map on top of the No. 7 Coal Member prepared by Clegg (Reference 57), Figure 2.5-285. The structural contour map on top of the No. 2 Coal Member correlates with structure contour maps on top of the No. 2 Coal Member (Reference 57) and on top of the Middle Devonian strata (Reference 58) except that the lower horizons are more intensely folded. No faulting has been recognized in association with the foregoing structural features either from aerial photographs, ERTS imagery, geophysical studies, borehole control, or excavation mapping. The glacial materials are devoid of lineaments or off-sets suggestive of faulting. Even if the bedrock unit elevation differences could be attributed to structural deformation, the relatively flat-lying and undeformed Pleistocene drift overlying bedrock demonstrates that the stresses which would have been responsible for the deformation have been inactive since at least pre-Pleistocene time. The Downs Anticline and its associated axial domes are stable and are of no structural significance at the site. 2.5.1.2.4 Site Surficial Geology The surficial geology is shown in Figure 2.5-15 and is discussed in Subsection 2.5.1.2.2.1. 2.5.1.2.5 Site Geologic History The geologic history of the site area is derived partly from exposures of soil units and partly from onsite borings in which samples of soil and bedrock were obtained. The deepest borings within the site area penetrated the uppermost bedrock formations of Pennsylvanian age. Discussions of events which occurred at the site during pre-Pennsylvanian times are based on data derived from adjacent regions. 2.5.1.2.5.1 Precambrian Era (Greater than approximately 600 million Years B.P.)

The Precambrian is the oldest recognized division of geologic time. The history of events which occurred during this long period of time is obscured by deep burial under younger rocks. No borings at or near the site have reached Precambrian rocks. Data from the regional area CPS/USAR CHAPTER 02 2.5-29 REV. 11, JANUARY 2005 suggest, however, that the Precambrian rocks in Illinois are igneous rocks, composed of granite, rhyolite, and associated rocks that formed in the interval from 1.1 to 1.4 billion years ago (Reference 2). The Precambrian basement in Illinois underwent a long period of erosion lasting from Late Precambrian time into Cambrian time (Reference 2). Consequently, the Precambrian surface is, in part, an erosional surface which may have several hundred feet of relief (Reference 2). The elevation of the Precambrian basement in the site vicinity is estimated to be approximately -6000 feet MSL (Reference 2) at a depth of approximately 6700 feet. 2.5.1.2.5.2 Paleozoic Era (Approximately 600 to 225 +/- 5 million years B.P.

The Clinton Power Station is located in the northern part of the Illinois Basin. This area was subject to intermittent tectonic movements occurring in the Illinois Basin throughout Paleozoic time. 2.5.1.2.5.2.1 Cambrian Period (Approximately 600 to 500 million years B.P.)

No onsite borings have penetrated Cambrian-age deposits. Indications are that the site area was submerged during Late Cambrian time. The first deposits in the advancing sea were coarse sand and fine pebbles, followed by finer sand, dolomite, and shale with an increasing amount of calcareous material. Before the close of Cambrian time, the seas cleared and chemical and/or organic precipitates which formed carbonate rocks were deposited. At the close of Cambrian time, the site area was uplifed. This was followed by a brief period of erosion (Reference 2). Approximately 3100 feet of Late Cambrian sediments underlie the site (Reference 2). 2.5.1.2.5.2.2 Ordovician Period (Approximately 500 to 430 +/- 10 million years B.P.)

No borings at the site have reached Ordovician age deposits. The Ordovician Period began with a transgression of the sea. General conditions favored the accumulation of calcereous deposits. At the close of Early Ordovician time, the sea again receded and a prolonged period of erosion was initiated. Later, the readvancing sea deposited a considerable quantity of fine to medium sand (the St. Peter Sandstone), followed by a thick sequence of calcareous deposits, and ending with accumulations of silt and clay

. Approximately 1000 feet of Ordovician sediments underlie the site (Reference 2). 2.5.1.2.5.2.3 Silurian Period (430 +/- 10 to 400 +/- 10 million years B.P.)

No borings at the site have reached Silurian-age deposits. However, much of the Illinois Basin, including the site area, was continuously beneath shallow seas during the Silurian Period.

Depositon in these shallow seas consisted primarily of carbonates, and reefs developed in some areas. Regional data suggest that carbonate deposition continued from Silurian into

Devonian time. Approximately 450 feet of Silurian age sediments underlie the site (Reference 2). 2.5.1.2.5.2.4 Devonian Period (400 +/- 10 to 340 +/- 10 million years B.P.)

No borings at the site have reached Devonian-age deposits. A major period of uplift and erosion occurred between the end of Lower Devonian sedimentation and the beginning of Middle Devonian sedimentation in the regional area, including the site vicinity. Sedimentation in shallow Middle Devonian seas in the site area probably consisted of limestones and dolomites CPS/USAR CHAPTER 02 2.5-30 REV. 11, JANUARY 2005 changing to black and gray shales in Late Devonian time. Total thickness of Devonian strata in the site area is about 200 feet (Reference 2). 2.5.1.2.5.2.5 Mississippian Period (340 +/- 10 to 320 +/- 10 million years B.P.)

No borings at the site have reached Mississippian age deposits. During Mississippian time, the deposition of fine clastics continued, although conditions gradually changed to alternating deposits of carbonates, silt, and sand. The close of Mississippian time was marked by uplift and erosion. Uplift of the Wapella East Dome portion of the Downs Anticline, near the site area (Subsection 2.5.1.2.3), took place during or prior to Mississippian time, and continued after the deposition of Mississippian age sediments (Reference 57), prior to Pennsylvanian deposition. 2.5.1.2.5.2.6 Pennsylvanian Period (320 +/- 10 to 270 +/- 5 million years B.P.)

The cyclical units of Pennsylvanian age strata indicate alternating periods of marine, nearshore, deltaic, and continental deposition which took place in much of the regional area throughout Pennsylvanian time. Nearly 600 feet of Pennsylvanian age strata were deposited in the site

area (Reference 2). Uplift on the Wapella East Dome portion of the Downs Anticline near the site occurred during and/or after Pennsylvanian time (Reference 57). A comparison of the relative amount of uplift occurring on the Downs Anticline before Pennsylvanian deposition and during and/or after Pennsylvanian deposition can be made by comparing the profiles of the tops of the Trenton (Ordovician) and Hunton (Silurian-Devonian) with the tops of the No. 2 and No. 7 Coal Members (Pennsylvanian) in Figures 2.5-283 and 2.5-286. 2.5.1.2.5.2.7 Permian Period (270 +/- 5 to 225 +/- 5 million years B.P.)

There are no deposits of Permian age in the regional area. It has been speculated that some Permian age deposits might have been deposited in the Illinois Basin (References 2 and 14), but if so, they were completely removed by subsequent erosion. 2.5.1.2.5.3 Mesozoic Era (225 +/- 5 to 65 +/- 2 million years B.P.)

2.5.1.2.5.3.1 Triassic Period (225 +/- 5 to 190 +/- 5 million years B.P.)

There are no deposits of Triassic age at the site. This was largely a period of erosion throughout the regional area (Reference 2). 2.5.1.2.5.3.2 Jurassic Period (190 +/- 5 to 135 +/- 5 million years B.P.)

There are no deposits of Jurassic age at the site. This was largely a period of erosion throughout the regional area (Reference 2). 2.5.1.2.5.3.3 Cretaceous Period (135 +/- 5 to 65 +/- 2 million years B.P.

) There are no Cretaceous age deposits at the site. This was probably a period of erosion in the site area.

CPS/USAR CHAPTER 02 2.5-31 REV. 11, JANUARY 2005 2.5.1.2.5.4 Cenozoic Era (65 +/- 2 million years B.P. to the present) 2.5.1.2.5.4.1 Tertiary Period (65 +/- 2 to 2 +/- 1 million years B.P.)

There are no Tertiary age deposits at the site. This was probably a period of erosion in the site area. 2.5.1.2.5.4.2 Quaternary Period (2 +/- 1 million years B.P. to the present)

Continental glaciers advanced over the site area several times during the Quaternary Period. Initial deposits in the site area are believed to be associated with the Kansan Stage (called pre-Illinoian Stage by Boellstorf, Reference 4), and consist of clean, fine to medium sand which accumulated in erosional valleys formed on the irregular Pennsylvanian bedrock surface.

These deposits are best illustrated in a subsurface section in Figure 2.5-279. The accumulations of sand in the bedrock valleys may have been the result of glaciation some distance upstream. The sand deposition was followed by a relatively thin accumulation of fine alluvial or lacustrine soils with some organic debris which also appears to be confined primarily to the bedrock valleys. Glacial ice of Kansan age eventually reached the site area depositing a blanket of glacial till over the entire area. At this time, some preexisting deposits were removed or reworked and the pre-existing erosional valleys were buried, forming a more or less featureless plain. The interglacial period following the final retreat of the Kansan glaciers is known as the Yarmouthian Stage. During this time the site area may have been covered by shallow lakes.

Topography was undoubtedly low with some vegetation in local areas. The site was reglaciated during the Illinoian Stage. A sequence of glacial tills, designated as the Glasford Formation, were deposited. Glaciolacustrine deposits and discontinuous lenses of sand and gravel may be interbedded within the glacial tills of the Glasford Formation. Following the final retreat of the Illinoian glaciers, the deposits of the Glasford Formation at the site were subjected to a long interval of weathering and erosion during and possibly after the interglacial interval known as the Sangamonian Stage. A thick soil (the Sangamon Soil) developed in the upper part of the deposits of the Glasford Formation, and these deposits are referred to as the weathered Glasford Formation in the FSAR and as Interglacial Zone, Sangamon Interglacial Zone, or Sangamon Soil Interval in the PSAR. During the Farmdalian Substage of the Wisconsinan Stage a thin organic silt, the Robein Silt, was deposited in the site area. Glaciers advanced over the site for the final time during the Woodfordian Substage of the Wisconsinan Stage, attaining their maximum extent shortly to the west of the CPS site.

Meltwater from the retreating Woodfordian ice sheet was responsible for erosion throughout the area. The existing Salt Creek probably was formed in this manner. Outwash and alluvial deposits in Salt Creek have probably accumulated since early Woodfordian time. Initially, runoff was relatively heavy and coarse-grained outwash deposits of the Henry Formation accumulated. As runoff decreased, fine-grained silts and clays of the Cahokia Alluvium accumulated along with some organic debris. The blanket of wind blown silt (loess) which covers the topographic highs of the site area was transported from sources along the major drainages of Illinois.

CPS/USAR CHAPTER 02 2.5-32 REV. 11, JANUARY 2005 2.5.1.2.6 Site Groundwater Conditions Groundwater conditions at the site are discussed in Subsection 2.4.13. 2.5.1.2.7 Geolocic Considerations Consideration has been given to all aspects of geology relevant to the suitability of the CPS site, including zones of soft or potentially liquefiable soils, karst, tectonic folding and faulting, slope stability, and the effects of man's activities at or in the vicinity of the site including surface or subsurface subsidence. The subgrades for all Category I structures are described in Subsection 2.5.4. There is no known karst development at the site or in the vicinity of the site. No evidence for tectonic faulting or Pleistocene or Holocene folding was noted in preliminary investigations or in the excavations for plant structures at the site. The stability of slopes at the site is discussed in Subsection 2.5.5.

There are no known instances of, or potential possibilities for, surface or subsurface subsidence, uplift, or collapse resulting from the activities of man within the site area. Present and former activities within the site area have included the removal of sand and gravel and the domestic use of groundwater. Sand and gravel production has been limited to surficial mining operations, and thus no hazard is posed to the plant site because of subsidence. There are no large uses of groundwater nor any industrial disposal wells in this area. No surface subsidence or response due to groundwater withdrawals have been reported near the site. Two oil fields are located within 15 miles of the Clinton Power Station (Figure 2.5-287). The Wapella East field is located approximately 6 miles northwest of the site, and the Parnell field is located approximately 7 miles northeast of the site (Reference 59). Both of these oil fields are located on domal structures along the Downs Anticline (Figure 2.5-285). There have been no instances of uplift, subsidence, or collapse associated with these oil fields, and no hazard is posed to the plant site because of these oil field developments. Five gas storage projects are located within 35 miles of the Clinton Power Station (Reference 60) (Figure 2.5-287). The Hudson gas storage project is located approximately 27 miles north of the site; the Lexington gas storage project is located approximately 30 miles north of the site, and the Lake Bloomington project is located approximately 34 miles north of the site (Figure 2.5-287). Each of these gas storage projects was developed by the Northern Illinois Gas Company, and for each of these projects, the storage reservoir is the Cambrian age Mt. Simon Sandstone

(Reference 60). The Manlove gas storage project is located approximately 23 miles east-northeast of the site (Figure 2.5-287). This gas storage project is operated by The Peoples Gas, Light, and Coke Company. The storage reservoir is the Cambrian age Mt. Simon Sandstone (Reference 60). The Lincoln gas storage project is located approximately 30 miles west of the site (Figure 2.5-287). This gas storage project is operated by the Central Illinois Light Company. The storage reservoir is in Silurian dolomite (Reference 60). There have been no instances of uplift, CPS/USAR CHAPTER 02 2.5-33 REV. 11, JANUARY 2005 subsidence, or collapse associated with these gas storage projects, and no hazard is anticipated to the plant site because of these gas storage projects. 2.5.2 Vibratory Ground Motion This section consists of a discussion and evaluation of the seismic and tectonic characteristics of the Clinton Power Station site and the surrounding region, and presents the rationale used to develop the seismic design criteria for the Clinton Power Station.

2.5.2.1 Seismicity 2.5.2.1.1 Seismicity Within 200 Miles of the Site The North Central United States is among one of the least seismically active areas of the United States. Since this area has been populated for almost 200 years, it is likely that most earthquake events of Intensity VI and all events of Intensity VII or larger on the Modified Mercalli (MM) Scale (Table 2.5-61) which have occurred during this time span have been reported. One hundred and sixty-four earthquakes greater than Intensity III are known to have occurred within 200 miles of the site. Table 2.5-4 lists all known reported events greater than Intensity III which have occurred between 37

° to 45° north latitude and 84

° to 93° west longitude and their locations are shown in Figure 2.5-288. The instrumentally determined locations are probably accurate to about

+/- 0.1°. The location of older events, not determined instrumentally, may have occurred as much as

+/- 0.5° from the stated location as the reported epicentral locations for these events normally correspond to the locations of the nearest reporting population center. There is no record of earthquakes with an Intensity of VIII or greater within 200 miles of the site; the closest Intensity VIII was 250 miles south at Charleston, Missouri, occurring in 1895 (Figure 2.5-426). If a shock of this size had occurred when the region was only sparsely settled, it almost certainly would have been mentioned in private journals or diaries, or preserved in Indian traditions, as had been the case in other regions. The lack of such documentation indicates the absence of significant earthquake activity within 200 miles of the site for a long period of time. The greatest earthquakes occurring within 200 miles of the site are listed below. These events had epicentral intensities of MM VII. a. May 26, 1909 - S. Beloit, Illinois; b. July 18, 1909 - central Illinois (Havana); c. September 27, 1909 - southeastern Illinois; and d. November 9, 1968 - southern Illinois. Isoseismal maps for these events are reproduced on Figures 2.5-289, 2.5-290, 2.5-291, and 2.5-292. These maps are useful since they show the geographic range of the areas of highest intensity. The greatest intensity induced at the Clinton site by these four earthquakes was MM-V. The closest occurrence of an epicentral intensity (MM) VI to the site was at a distance of approximately 100 miles. Therefore, the maximum intensity experienced at the Clinton site from any earthquake occurring within a 200-mile radius of the site was MM-V.

CPS/USAR CHAPTER 02 2.5-34 REV. 11, JANUARY 2005 2.5.2.1.2 Distant Events 2.5.2.1.2.1 Central Stable Region Only one earthquake occurring in the Central Stable Region at distances more than 200 miles from the site has been felt at the site itself. Docekal (Reference 61, Plate 3) shows Intensity I-III (MM) at the site from the March 8, 1937 Anna, Ohio earthquake (Subsection 2.5.2.3.1.11). 2.5.2.1.2.2 Mississippi Embayment Area The largest recorded earthquakes which have occurred in the central part of the United States were the New Madrid events of 1811-1812. These events occurred in the Mississippi Embayment area of the Gulf Coast Tectonic Province (References 62, 63, and 64) at a distance of over 250 miles from the site (Figure 2.5-293 and Table 2.5-5). Over a period of 3 months during 1811-1812, at least 250 minor events and three major separate shocks occurred, the largest of which had an Intensity (MM) of XI-XII (References 65 and 66). There has been no recurrence of such a major earthquake in this zone, but there is evidence of activity prior to the New Madrid events. There is a report of a very large shock on December 25, 1699, with its epicenter in western Tennessee, which shook approximately the same area as the 1811-1812 events. Written records also indicate that "notable vigorous" shocks occurred in 1776, 1791 or 1792, 1795, 1796, and 1804. Indian traditions also record a previous earthquake which devastated the same area (Reference 65). In addition to these events, an Intensity (MM) VIII event occurred in 1895 in Charleston, Missouri also within the Mississippi Embayment area, which was probably felt at the Clinton site approximately 250 miles away.

2.5.2.1.2.3 Other Events One other event was probably felt at the Clinton Power Station site: the 1886 Intensity (MM) X Charleston, South Carolina event which occurred in the Atlantic Coastal Province. Details of these and other distant events are presented in Table 2.5-5. See Dutton, C. E. (1888), The Charleston Earthquake, U. S. Geological Survey 9th Annual Report, Pages 203-528. 2.5.2.2 Geologic Structures and Tectonic Activity The Clinton site and the vast majority of the 200-mile radius site region lie within the Central Stable Region of the North American Continent (Reference 63). This region is characterized by a relatively thin veneer of sedimentary rocks overlying a crystalline basement. These areas were deformed principally by movements which occurred as a result of tectonic activity during the Paleozoic resulting in a series of gentle basins, domes, and other structures. Since the end of the Paleozoic, the area has remained generally quiescent. A few square miles of the southernmost area of the site region overlaps the Mississippi Embayment region of the Gulf Coastal Plain Tectonic Province (Reference 63). The site is located within the Illinois Basin. The most significant nearby structure is the Downs anticline which is genetically related to the La Salle Anticlinal Belt. A description of the faulting and tectonic features in the area is presented in Subsections 2.5.1.1, 2.5.1.2, and 2.5.3.

CPS/USAR CHAPTER 02 2.5-35 REV. 11, JANUARY 2005 2.5.2.3 Correlation of Earthquake Activity with Geologic Structures or Tectonic Provinces The Central Stable Region Tectonic Province is generally noted for its lack of significant seismic activity. To evaluate the earthquake potential of the Clinton site, two different approaches were utilized to correlate earthquake activity with geologic structures and/or tectonic provinces: (1) the 200-mile radius site region was subdivided into seismotectonic regions (Subsection 2.5.2.3.1) utilizing methods similar to Stearns and Wilson (Reference 67), and (2) analysis of the relationship of the site to zones of relatively high seismicity within the Central Stable Region Tectonic Province and the Gulf Coast Plain Tectonic Province (Subsection 2.5.2.3.2) were carried out. 2.5.2.3.1 Seismogenic Regions Eleven seismogenic regions can be delineated within 200 miles of the Clinton Power Station, primarily on the basis of structure. These subdivisions also indicate the differing geologic and seismic histories of the seismogenic regions. The following subsections describe the eleven seismogenic regions within the 200-mile radius site area and other regions pertinent to the site. Each region is outlined in Figure 2.5-288.

2.5.2.3.1.1 Illinois Basin Seismogenic Region The site is located in the center of the Illinois Basin Seismogenic Region. The north and northeastern boundaries of this region correspond to and are defined by the limits of the Plum River and Sandwich Fault zones. The region has experienced 60 recorded earthquakes, the largest of which was Intensity (MM) VII. A tentative correlation of some events (with areas of steep gradients in the earth's gravitational field) has been proposed by various authors, notably McGinnis and Ervin (Reference 68). They interpret the gradients as boundaries separating crustal blocks of different densities. However, based on the present state of knowledge, these events are considered random. Therefore, the possibility of an Intensity (MM) VII event anywhere in the basin must be considered. 2.5.2.3.1.2 Ste. Genevieve Seismogenic Region The Ste. Genevieve Region lies approximately 175 miles southwest of the site and is related to and defined by the imbricated Ste. Genevieve Fault System. This region exhibits a characteristic maximum intensity earthquake of (MM) VI. While there is no geological evidence of capable faulting in this region, fault plane solutions coincide with the trace of the Ste.

Genevieve fault. The boundary with the Illinois Basin is based both on a change in structure and by a contrast in seismicity. See Street, R. L. et al., 1974, Earthquake Mechanisms in the Central U. S., Science, V. 184, Pages 1285-1287. 2.5.2.3.1.3 St. Francois Mountains Seismogenic Region This seismogenic region is defined by the limits of a region of Precambrian rock outcrops in southeast Missouri, constituting the exposed core of the Ozark Uplift. Moderate earthquake activity is associated with faults on the margin of the St. Francois Mountains Regions. The maximum seismic events recorded are Intensity (MM) VI-VII. The closest distance between this region and the site is approximately 175 miles.

CPS/USAR CHAPTER 02 2.5-36 REV. 11, JANUARY 2005 2.5.2.3.1.4 Chester-Dupo Seismogenic Region The Chester-Dupo Region is defined by an area of faulting and folding in the vicinity of St. Louis (Reference 71). This region, approximately 110 miles southwest of the site, is one of moderate seismicity with maximum events characteristic of (MM) VII. The boundary between the Chester-Dupo Seismogenic Region and the Illinois Basin Seismogenic Region is marked by the transition from the folds and faults to the deeper, structurally less complex Illinois Basin. This region marks a hinge line between the Illinois Basin and the northeast boundary of the Ozark

Uplift. 2.5.2.3.1.5 Wabash Valley Seismogenic Region This seismogenic region is defined by the limits of the Fairfield Basin, the deepest part of the Illinois Basin, and by the northwest-trending faults of the Wabash Valley. The closest approach of this region to the site is approximately 70 miles This area has moderate seismicity with maximum events of the (MM) VII. Events in this region occur more frequently than events in the adjoining parts of the Illinois Basin (Reference 62). The boundaries of the Wabash Valley Seismogenic Region are well defined by structure and geological history in addition to its seismic pattern. 2.5.2.3.1.6 Western Kentucky Fault Zone Seismogenic Region This region, 175 miles from the Clinton site, consists of a fault system bending north 80

° east. The western boundary with the Illinois Basin Seismogenic Region and the Ste. Genevieve Seismogenic region is defined by a change in seismicity. The southern boundary of the region lies along the boundaries of the New Madrid Region and East Mississippi Embayment Regions and along the northern flank of the Tennessee-Kentucky Stable Region. The Western Kentucky Fault Zone is a stable area with only a few randomly occurring epicenters (Reference 67). 2.5.2.3.1.7 Iowa-Minnesota Stable Seismogenic Region This region is one of extremely low seismicity with a general maximum intensity of (MM) V. The boundary between this region and the Illinois Basin is approximately 120 miles from the site and is marked by a gentle zone of flexure known as the Mississippi River Arch. 2.5.2.3.1.8 Missouri Random Seismogenic Region The Missouri Random Seismogenic Region is bounded by the Chester-Dupo Region to the east, and its contact with the Illinois Basin Region is marked by the Lincoln Fold. This region lies approximately 100 miles southwest of the site. This area is characterized by the occurrence of random seismic events of maximum (MM) V which are not associated with any known structure. 2.5.2.3.1.9 Michigan Basin Seismogenic Region The Michigan Basin Seismogenic Region is an area of extremely low seismicity with a total of 10 recorded events. The largest seismic event was an (MM) VI. This area is separated from the Illinois Basin by the Kankakee Arch and lies approximately 150 miles northeast of the site.

CPS/USAR CHAPTER 02 2.5-37 REV. 11, JANUARY 2005 2.5.2.3.1.10 Eastern Interior Arch System Seismogenic Region This region is composed of a series of gentle Paleozoic arches and domes within the eastern part of the Central Stable Region. Structurally, this area is composed of the Wisconsin, Kankakee, Findlay, and Cincinnati Arches, Jessamine Dome, and Wisconsin Dome. While this system can be subdivided into the various structures, the geological history of the structures and lithologies as well as general patterns of seismicity are similar. Inasmuch as the boundaries between any of the structures are rather nebulous, divisions would be rather

arbitrary. The Wisconsin Dome in the northern part of the Central Stable Region consists of Precambrian rocks and bears more similarity to the Laurentian Shield subdivision of the Central Stable Region of Canada than to the Interior Lowlands subdivision within the United States (References 63 and 64). The Wisconsin Dome is an extremely stable part of the Central Stable Region and represents the most seismically stable part of this region with maximum seismic activity of (MM) V. The Wisconsin Arch is defined structurally by the low, north south trending, uplifted area extending south from the Wisconsin Dome and is herein defined to include the east-west trending cross-cutting folds and faults of southern Wisconsin. The boundary between the Wisconsin Arch and the Kankakee Arch is extremely hard to define. The name changes from the Wisconsin Arch to the Kankakee Arch northwest of Kankakee, Illinois. The Wisconsin Dome and Arch have a Precambrian core and are believed to have acquired their relief primarily by uplift whereas the relief on the Kankakee Arch is due primarily to more rapid subsidence on the bordering Michigan and Illinois Basins. The arch system continues southeastward to join the Cincinnati Arch and the Jessamine Dome. The Findlay Arch is a northeastward splay off the Cincinnati Arch and separates the Michigan Basin from the Appalachian Basin. Seismicity within this region is generally of (MM) V. However, isolated events of (MM) VII have occurred which cannot be related to specific structures. Therefore, the entire region must be assigned a maximum potential random event of (MM) VII. 2.5.2.3.1.11 Anna Seismogenic Region The Anna Region lies at the intersection of the Kankakee, Findlay, and Cincinnati arches in western Ohio, 220 miles east of the site. This area has experienced continued and moderately severe seismic activity. The largest historic earthquakes commonly have been of Intensity (MM)

VII, with a single event on March 8, 1937, which has been assigned a maximum Intensity (MM)

VII-VIII by Coffman and Von Hake (Reference 72). However, a detailed analysis indicates that this event should be reclassified as maximum Intensity (MM) VII (Reference 69). The Anna Region is defined as lying within a basement structural zone bounded on the south by a northwest-trending band of basement faulting, on the east by a zone of structural weakness marked by a north-south trending band of magnetic highs and lows, on the north by a change from igneous extrusive to igneous intrusive rock, and on the west by the change from acidic extrusive to basic extrusive rocks (Reference 69 and 70). The combination of geological features within this area is unique. There is no other known area within the central United States with the combination of factors similar to this region. The earthquake events which have CPS/USAR CHAPTER 02 2.5-38 REV. 11, JANUARY 2005 occurred in this region are not random but rather the result of the unique combination of geological phenomena (Reference 69). 2.5.2.3.1.12 New Madrid Seismogenic Region The New Madrid Seismogenic Region can be defined approximately on any tectonic map as corresponding to the northern portion of the Mississippi Embayment which is the northern portion of the Gulf Coastal Plain Tectonic Province (References 62, 63, and 64, and Figures 2.5-288 and 2.5-294). The New Madrid events of 1811-1812 were the largest earthquakes ever experienced in the central and eastern United States. Chimneys were dislodged as far north as St. Louis, Missouri and the aftershocks from these events continued for 2 years (Reference 73). These events occurred more than 260 miles from the Clinton site. Extensive studies have been conducted to determine the northernmost region in which these events could occur. This has been documented in a Sargent & Lundy report, dated May 23, 1975, entitled, "Supplemental Discussion Concerning the Limit of the Northern Extent of Large Intensity Earthquakes Similar to the New Madrid Events" (Reference 74). Further discussion on this matter took place at a meeting held on January 26, 1976, in the offices of the Illinois State Geological Survey, Urbana, Illinois. Representatives were present from the Nuclear Regulatory Commission, the Illinois State Geological Survey, the Indiana Geological Survey, the Kentucky Geological Survey, St. Louis University, Sargent & Lundy, Dames & Moore, and Seismograph Service Corporation (Birdwell Division). The scientific data presented clearly indicated that the New Madrid area, at the intersection of the Pascola Arch and the Ozark Dome, is tectonically unique and that the northernmost extent of the structurally complex New Madrid area is conservatively taken as 37.3

° north, 89.2

° west - 200 miles from the site. It remains the applicant's interpretation, based on tectonic, geophysical and seismic data, that New Madrid-type events should not extend across Tectonic Province boundaries and up the Wabash Valley Fault System. These conclusions were also presented to the NRC (formerly the AEC) previously at: a. AEC staff review meeting in Bethesda, Maryland, for the Clinton Power Station, on June 17, 1974. b. ACRS Subcommittee Hearings in Urbana, Illinois, for the Clinton Power Station, on March 29, 1975. c. ACRS Subcommittee Hearing in Bethesda, Maryland, for the Clinton Power Station, on April 4, 1975. d. ACRS Subcommittee Hearing in Madison, Indiana, for the Marble Hill Nuclear Generating Station, on October 1, 1976. e. ACRS Full Committee meeting in Washington, D.C., for the Marble Hill Nuclear Generating Station, on October 14, 1976. A regional microearthquake detection network has recently been installed in this area. Analysis of data obtained from this network indicates that the New Madrid Region and the Wabash Valley Region are probably two distinct seismic regions (Reference 75).

CPS/USAR CHAPTER 02 2.5-39 REV. 11, JANUARY 2005 While all evidence indicates that a New Madrid-type event could only occur in the area of the Pascola Arch, if the events are transposed northward, the major crustal discontinuity along the Rough Creek Fault Zone serves as a boundary for further northward migration. Even if this zone is selected as the northern boundary, the New Madrid-type events could occur no closer than 170 miles from the site. 2.5.2.3.2 Tectonic Provinces 2.5.2.3.2.1 Central Stable Region Tectonic Province The Central Stable Region is noted for its general lack of significant seismic activity with the largest events generally of (MM) VII. Within this tectonic province there are several zones of relatively high activity. These are (1) near Attica, New York, (2) near Anna, Ohio, (3) the Wabash River Valley of southern Illinois and Indiana, (4) in eastern Kansas and Nebraska along the midcontinent gravity and magnetic high in the area of the Nemaha Anticline, and (5) near St.

Louis, Missouri (Figure 2.5-294). The Attica events are associated with the Clarindon-Lindon Structure and the August 12, 1929 event has been assigned an Intensity (MM) VIII by Coffman and von Hake. However, the amount of damage and estimated magnitude of this event indicate that it should be reclassified in the record as an Intensity VII (Reference 76). The area around Anna, Ohio has experienced a relatively large amount of seismic activity compared to other areas of the Central Stable Region. As described in Subsection 2.5.2.3.1.11 the area of earthquake activity corresponds to a highly complex Precambrian structural zone. In addition, the March 8, 1937 event, which has been assigned an Intensity (MM) VII-VIII by Coffman and von Hake (Reference 72), has been analyzed and all indications are that this event has a maximum epicentral Intensity (MM) VII (Reference 69). The Wabash Valley Fault Zone was described in Subsection 2.5.2.3.1.5 and has had maximum recorded seismic activity of Intensity (MM) VII. Several events of Intensity (MM) VII have occurred in the area of the Nemaha Anticline magnetic high. The relationship of earthquake activity to the midcontinent gravity and magnetic high has been documented in Subsection 2.5.2 of the Wolf Creek PSAR (Reference 77). The activity near St. Louis, Missouri has been assigned to the Chester-Dupo Region documented in the PSAR for the Callaway Plant (Reference 71). Historical activity in this area has attained a maximum of Intensity (MM) VII. In addition to these areas of the Central Stable Region which have had relatively high seismic activity, an Intensity (MM) VIII event was reported in the Keewenaw Peninsula of Michigan in 1906 (Reference 72). The felt area of the 1906 event was approximately equal to that of an average Intensity (MM) III-IV event (Reference 72). The area of the epicenter is highly faulted, and the areas of damage and perceptibility coincide with areas of mining activity. Smaller events which occurred earlier in the year as well as the larger event in 1906 all appear directly attributable to mining activity (Reference 61).

CPS/USAR CHAPTER 02 2.5-40 REV. 11, JANUARY 2005 2.5.2.3.2.2 Gulf Coastal Plain Tectonic Province The New Madrid events of 1811-1812 occurred in the Gulf Coastal Plain Tectonic Province, not in the Central Stable Region Tectonic Province. These events are associated with a highly complex structural zone near the crest of the Pascola Arch (Subsection 2.5.2.3.1.12). If these events are translated to the closest approach of this tectonic province to the site, these events could be expected to occur no closer than 195 miles from the site or 65 miles closer to the site than the 1811-1812 events occurred. 2.5.2.3.3 Earthquake Events Significant to the Site From both types of analysis of the association of earthquakes with structure as described above, the most significant earthquakes in the region are the July 18, 1909, Intensity (MM) VII central Illinois (Havana) earthquake, the May 26, 1909, Intensity (MM) VII northern Illinois earthquake, the September 27, 1909, Intensity (MM) VII southeastern Illinois - southwest Indiana earthquake, the 1968 Intensity (MM) VII southern Illinois earthquake, and the New Madrid earthquakes of 1811-1812. This evaluation is based on epicentral intensity, distance from the site and tectonic association. 2.5.2.4 Maximum Earthquake Potential Based on the discussion in Subsection 2.5.2.3, the maximum earthquake which could be expected would be a repetition of the 1909 Havana Intensity (MM) VII event near the site. This is equivalent also to the occurrence of the largest event which has ever been recorded within the Central Stable Region, and which cannot yet be associated with a specific structure or structural region; it is, therefore, described as random. The level of ground motion experienced from a near field Intensity (MM) VII event would be expected to envelope the motion from a recurrence of a New Madrid-type event at the closest approach of the Mississippi Embayment, a distanc e of approximately 195 miles from the site (Subsection 2.5.2.6). 2.5.2.5 Seismic Wave Transmission Characteristics of the Site The engineering properties of the soils and bedrock units at the site were evaluated using field geophysical measurements and laboratory testing; the properties determined by laboratory testing are discussed in Subsection 2.5.4.2.2. Geophysical investigations performed at the plant site are presented in Subsection 2.5.4.4. The velocity of compressional and surface wave propagation and other dynamic properties of the natural subsurface conditions were evaluated from these investigations and the data were used in analyzing the response of the materials to earthquake loading. Dynamic moduli for the subsurface soil and rock at the site were calculated based on measured properties. The in situ field measurements were compared with laboratory tests on the same materials. These analyses are presented in Subsection 2.5.4.7. These data were used in studies of the site dynamic response.

CPS/USAR CHAPTER 02 2.5-41 REV. 11, JANUARY 2005 2.5.2.6 Safe Shutdown Earthquake The recommended safe shutdown earthquake (SSE) was defined as the occurrence of an Intensity (MM) VII event near the site. This near field event can be correlated to a mean horizontal ground acceleration of 0.13g (Reference 78 and Figure 2.5-295). This level of ground motion would be expected to envelope the motion from a recurrence of a New Madrid-type event at the closest approach of the Mississippi Embayment at a distance of 195 miles from the site. At the time of the review of the construction permit application, the NRC staff insisted that the safe shutdown earthquake be defined as an Intensity (MM) VIII event near the site. The Illinois Power Company considers this position to be extremely conservative and inconsistent with the seismicity and tectonics of the site region.

However, in order to expedite licensing, the NRC staff position was adopted. This resulted in a maximum horizontal ground surface acceleration at the site of 0.25g. To provide an additional margin of safety, this value was applied at foundation level in the free field. Utilizing the subsurface properties presented in Subsection 2.5.4.7, the corresponding ground surface acceleration was found to be 0.26g. The NRC staff also took the position that the 1811-1812 New Madrid-type earthquakes must be considered to occur at 110 miles from the site, near Vincennes, Indiana. In order to expedite licensing, the effects of the occurrence of such an event was evaluated, and it was shown that the motions generated by the distant event were enveloped by the motions caused by the near field earthquake. The free field ground response spectra prepared in accordance with Regulatory Guide 1.60 for a horizontal ground acceleration of 0.26g are presented in Figure 2.5-296. SSER 3 concluded that the site specific spectra are roughly equivalent to 0.20g anchored to the Reg Guide 1.60 spectra is acceptable. 2.5.2.7 Operating Basis Earthquake The operating basis earthquake (OBE) is intended to indicate those levels of ground motion which could reasonably be expected to occur at the plant during the operating life of the facility.

As such, the OBE has been selected on the basis of a seismic risk analysis, the results of which are presented in Subsection 2.5.2.7.1. 2.5.2.7.1 Seismic Risk Analysis A quantitative seismic risk analysis of the site has been carried out using the method developed by Merz and Cornell (Reference 79 (a)). The output of this analysis is a plot of annual risk (probability of exceeding acceleration) versus peak ground acceleration at the site. The risk analysis is based on the historic seismicity in each of the seismotectonic provinces described in Subsection 2.5.2.3. The following earthquake sources have been considered in the risk analysis: a. Illinois Basin Seismogenic Region,

b. Ste. Genevieve Seismogenic Region, CPS/USAR CHAPTER 02 2.5-42 REV. 11, JANUARY 2005 c. St. Francois Mountains Seismogenic Region, d. Chester-Dupo Seismogenic Region,

e. Wabash Valley Seismogenic Region,

f. Western Kentucky Fault Zone Seismotectonic Region, g. Iowa-Minnesota Stable Seismogenic Region, h. Missouri Random Seismogenic Region,

i. Michigan Basin Seismogenic Region,

j. Eastern Interior Arch System Seismogenic Region,

k. Anna Seismogenic Region, and

l. New Madrid Seismogenic Region. Other seismotectonic provinces discussed in Subsection 2.5.2.3 are either too far from the site or have seismicity too low to require consideration.

The mean annual number of events ( MM Intensity IV), known as "activity rates," for the various sources are given in Table 2.5-65. These rates have been obtained by dividing each source's total number of events between 1879 and 1978 by 100. The table also shows the historical maximum Mercalli intensity for each source zone, as discussed in Subsection 2.5.2.3. The probability distribution of events of different epicentral intensities, I e, is chosen in the form recommended in Reference 79(b):

ln {P [Ie ie]} = -1(ie - i ) for i < ie < iu (2.5-1) where i and iu are the lower and upper epicentral intensity bounds and 1 is a shape coefficient which is conservatively estimated as 0.863 (Reference 80). The lower bound intensity for the local earthquake source used in the analysis is (MM) IV; for all other sources the value is (MM)

V. The upper bound intensity used in the analysis for each source is shown in Table 2.5-65. The spatial attenuation of intensity (Equation 2.5-2) in the central United States as developed by Gupta (Reference 81) is considered appropriate for the site region and is used in this analysis where Is is the site intensity and R is the epicentral distance in kilometers.

Is = Ie + 2.35 - 1.79 log 10R - 0.00316R (2.5-2) for R 20 km. The standard deviation, s, of the error in I s of this relationship is estimated as 0.45 by the same author. The correlation of site intensity, I s, and the peak ground acceleration, a, is obtained from Reference 82:

logl0a = 0.24I s + 0.26 (2.5-3) and the antilog of the standard error of a is 2.19.

CPS/USAR CHAPTER 02 2.5-43 REV. 11, JANUARY 2005 For the sources closer than 20 km to the site, the form of Equation (2.5-3) remains as it is since no attenuation of intensity (i.e., I s = Ie) is used for these sources in this study. For sources farther than 20 km from the site, the acceleration, a, depends on the epicentral intensity and the distance (i.e., by combining Equations (2.5-2) and 2.5-3)):

logl0a = 0.24I e + 0.824 - 0.0007584R - 0.4296log l0R (2.5-4) The computer program SRA (Seismic Risk Analysis) developed by MIT and updated by Sargent & Lundy has been employed in this analysis. The concept of risk calculations by SRA can be understood using the following example. The total number of events per year which will cause a site peak ground acceleration greater than or equal to 0.05g is the sum of contributions from individual sources. The latter are the sums of the expected numbers of events from the elemental areas making up the source. The expected number of events from each elemental area of the source is simply the product of (a) the mean annual activity rate, and (b) the fraction of events which will be large enough to generate a peak ground acceleration 0.05g at the site. How large such an event must be (i.e., how large its epicentral intensity must be to cause a peak ground acceleration 0.05g at the site) can be estimated by using the attenuation and intensity-peak ground acceleration relationships. The numerical procedure adopted herein (Reference 79) permits an accounting for the quoted uncertainties in these relationships. Once the required epicentral intensity is known, the fraction of all events that are equal to or greater than that intensity value can be obtained from the epicentral intensity probability distribution. The seismic risk curve due to all relevant seismic sources is shown in Figure 2.5-427. It shows the resulting annual risk that any specified peak horizontal ground acceleration will be exceeded. The operating basis earthquake for the plant has been selected as 0.10g applied at the foundation level. Utilizing the subsurface properties presented in Subsection 2.5.4.7, the corresponding ground surface acceleration was determined to be 0.11g. From Figure 2.5-427, it can be seen that the annual probability of exceeding this OBE acceleration at the site is 1.5xl0-3 per year (or mean recurrence interval of about 650 years). Based on this low probability, it is concluded that the selection of the OBE has been conservative. The free field ground response spectra prepared in accordance with Regulatory Guide 1.60 for a horizontal ground surface acceleration of 0.11g are presented in Figure 2.5-297. 2.5.3 Surface Faulting There is no evidence for surface faulting at the site or the area surrounding the site (200-mile radius around the plant site). Further, faults which have been mapped in Illinois have shown no sign of movement during Quaternary time. Based on the data contained in Subsections 2.5.1 and 2.5.2, and the interpretation and conclusions from those data, there are no capable faults, as defined by 10 CFR 100 Appendix A, within 200 miles of the site. The closest proposed fault to the site is the so-called Tuscola Fault (Reference 47), postulated to trend north-south, approximately 20 miles east of the Clinton Power Station. The existence of this fault has not been accepted by the Illinois State Geological Survey (Reference 48). The nearest confirmed fault is the Sandwich Fault Zone, located approximately 90 miles northeast of CPS/USAR CHAPTER 02 2.5-44 REV. 11, JANUARY 2005 the Clinton Power Station. The last movement on the Sandwich Fault Zone occurred during the interval from Post-Silurian to Pre-Pleistocene time, probably in the late Paleozoic (Reference 22). 2.5.3.1 Geologic Conditions of the Site A discussion of the lithologic, stratigraphic, and structural conditions of the site and the area surrounding the site, including its geologic history, is contained in Subsection 2.5.1. 2.5.3.2 Evidence of Fault Offset There is no evidence of fault offset at or near the ground surface at or near the site. The structural geology at the site and surrounding region is discussed in Subsections 2.5.1.1.4 and

2.5.1.2.3. 2.5.3.3 Earthquakes Associated with Capable Faults There have been no historically reported earthquakes within 5 miles of the site. No capable faulting is known to exist within 200 miles of the site. 2.5.3.4 Investigation of Capable Faults No capable faulting is known to exist within 200 miles of the site.

2.5.3.5 Correlation of Epicenters with Capable Faults No capable faulting is known to exist within 200 miles of the site, and no earthquake epicenter is known within 5 miles of the site. 2.5.3.6 Description of Capable Faults No capable faulting is known to exist within 200 miles of the site.

2.5.3.7 Zone Requiring Detailed Faulting Investigation Geologic investigations of the site have not found any evidence of capable faulting; therefore, the detailed fault investigation required for capable faulting is not needed. 2.5.3.8 Results of Faulting Investigation Geologic investigations of the site and the area surrounding the site have indicated that no capable faulting is present within 200 miles of the site, and that no surface faulting is present within 5 miles of the site; therefore, a study of surface faulting is not required. 2.5.4 Stability of Subsurface Materials and Foundations This section presents the evaluation of the stability of subsurface materials that underlie the site. This evaluation is based upon preliminary soils studies conducted prior to construction operations and the properties of the soils encountered during the earthwork construction.

CPS/USAR CHAPTER 02 2.5-45 REV. 11, JANUARY 2005 2.5.4.1 Geologic Features A detailed discussion of the geologic characteristics of the site is given in Subsection 2.5.1.2. 2.5.4.2 Properties of Subsurface Materials This subsection presents static and dynamic engineering properties of the subsurface materials encountered in the borings drilled as part of the investigation program, the properties of the materials used during the site excavation, and the properties of the materials used as structural fill and backfill. The material properties were based upon: a. a review of all available field and laboratory tests performed during this investigation, b. a review of the geophysical surveys performed during this investigation, c. a review of the latest available literature, d. a review of similar studies performed recently for nuclear generating stations at other locations, and e. testing performed during construction. Representative soil samples and rock cores extracted from the test borings were subjected to laboratory tests to evaluate the physical and chemical characteristics of the soil and rock encountered at the site. The physical and chemical characteristics of the groundwater encountered at the site were evaluated by testing representative samples obtained from test borings and piezometers at the site. The laboratory program for the PSAR stage included the following tests: a. strength test - 1. direct shear, 2. unconfined compression on undisturbed samples, 3. unconfined comppression on remolded samples,

4. unconsolidated-undrained triaxial compression, and

5. consolidated-undrained triaxial compression - some with pore pressure measurements; b. dynamic tests - 1. cyclic triaxial,

2. resonant column, and

3. shockscope; c. other physical tests -

CPS/USAR CHAPTER 02 2.5-46 REV. 11, JANUARY 2005 1. Atterberg limit, 2. consolidation,

3. in situ moisture and density determinations,

4. permeability, and 5. relative density; d. chemical tests on groundwater samples; and e. chemical analysis on coal samples. The results of the unconfined compression, unconsolidated-undrained triaxial, consolidated-undrained triaxial, moisture and density, and Atterberg limit tests are presented to the left of the logs of borings. Tests performed and reported elsewhere are indicated by symbols in the left-hand column of the logs of borings. The key to the test symbols is presented at the bottom of Figure 2.5-298. Subsection 2.5.4.2.6 presents soil investigations and testing programs undertaken, following the submittal of the PSAR, on in situ material as well as fill material under the guidance of Sargent

& Lundy personnel as well as the Quality Control Program instituted at the CPS site during the construction of the power station. These programs were initiated to confirm soil properties determined in initial investigations and used in the design considerations for the earthwork construction. 2.5.4.2.1 Strength Tests 2.5.4.2.1.1 Strength Tests on Soil Selected representative soil samples were tested in the manner described in Figure 2.5-299 to determine their strength characteristics. Unconsolidated-undrained triaxial compression and unconfined compression tests were performed on selected, undisturbed samples to evaluate their undrained strength characteristics. Consolidated-undrained triaxial, with and without pore pressure measurements, and direct shear tests were performed on samples to evaluate the effective stress parameters.

The samples tested were sheared under a surcharge or confining pressure corresponding approximately to the effective soil and/or surface loads occurring, or expected to occur, in the field. The triaxial and unconfined compression tests were run in the manner shown in Figure 2.5-299. The direct shear tests were run in the manner shown in Figure 2.5-300. A load-deflection curve was plotted for each test and the strength of the soil was determined from this curve. Determinations of the field moisture content and dry density of the soil were made in conjunction with each strength test. The results of the strength tests and the corresponding moisture content and dry density determination are presented to the left of the boring logs (Figures 2.5-19 through 2.5-270) and in Tables 2.5-6 through 2.5-17.

CPS/USAR CHAPTER 02 2.5-47 REV. 11, JANUARY 2005 2.5.4.2.1.2 Strength Tests on Rock The strengths of the underlying rock formations were evaluated by subjecting representative rock core sampling to unconfined compression tests (ASTM D2938-71). The tests were performed by subjecting samples approximately 4 inches in height and 2 inches in diameter to a constant rate of axial load. The modulus of elasticity of the rock was evaluated by recording the deformation of the rock and computing the stress-strain relationship. The results of the rock compression tests are presented to the left of the boring logs. 2.5.4.2.2 Dynamic Tests 2.5.4.2.2.1 Cyclic Triaxial Tests The behavior of representative soils under dynamic loading was evaluated by conducting dynamic triaxial tests in the manner indicated in Figure 2.5-301. The samples were initially allowed to consolidate under confining pressures representative of existing in situ conditions. Strain controlled dynamic triaxial tests were performed on the consolidated samples by subjecting them to sinusoidally varying axial strains at a frequency of 1 hertz. From these tests dynamic material properties such as strain dependent shear modulus and damping values were obtained. The test results are presented in Tables 2.5-18 through 2.5-25. Dynamic triaxial tests were performed on remolded samples of material similar to that used as structural fill beneath the plant and the test results are shown in Figures 2.5-302 through 2.5-311. Dynamic triaxial tests were also performed on representative in situ soils obtained from borings in the ultimate heat sink area. The results of these tests are shown in Figures 2.5-312 through 2.5-317 and in Table 2.5-22. 2.5.4.2.2.2 Resonant Column Tests Dynamic torsional shear (resonant column) test s were performed on representative soil and rock samples to evaluate the modulus or rigidity (shear modulus) of these materials in the manner described in Figure 2.5-318. The tests were conducted at natural moisture content over a range of confining pressures. The results of the resonant column tests are presented in Tables 2.5-26 through 2.5-29. 2.5.4.2.2.3 Shockscope Tests Compressional wave velocity (shockscope) tests were performed on representative soil and rock samples. The velocity observed in the laboratory was used to correlate with field velocity measurements obtained during the geophysical survey. In this test, samples are subjected to a physical shock, and the time required for the shock wave to travel the length of the sample is measured. The velocity of compressional wave propagation is then computed. The samples were tested in an unconfined state. The test results are presented in Table 2.5-30.

CPS/USAR CHAPTER 02 2.5-48 REV. 11, JANUARY 2005 2.5.4.2.3 Other Physical Tests 2.5.4.2.3.1 Atterberg Limit Tests Representative soil samples were tested to evaluate their plasticity characteristics (ASTM D424-59 and ASTM D423-66). The results of these tests were used for classification and correlation purposes. The Atterberg limit determinations are presented to the left of the boring logs. 2.5.4.2.3.2 Consolidation Tests Consolidation tests were performed on representative cohesive soil samples and granular borrow materials to determine compressibility characteristics of the soils. The tests were run in the manner described in Figure 2.5-319. Tests performed on Pitcher samples and the 4-inch core samples (high recovery core barrel method of sampling) were run using a procedure suggested by Bjerrum (Reference 83). The test results are presented in Figures 2.5-320 through 2.5-346. 2.5.4.2.3.3 In Situ Moisture and Density Determinations In addition to the in situ moisture and density determinations made in conjunction with the previous tests, additional moisture and density tests were performed on other soil samples for correlation purposes. The results of all moisture and density determinations are presented to the left of the boring logs. The moisture content determinations were performed according to ASTM D2216-66. 2.5.4.2.3.4 Laboratory Permeability Te sts Falling-head and constant-head type permeability tests were performed on representative soil samples to provide data for determining groundwater movements in the vicinity of the CPS site as well as in the vicinity of the main dam. The tests were performed in the manner described in Figure 2.5-347. The results of these tests are presented in Tables 2.5-31, 2.5-32, and 2.5-33. 2.5.4.2.3.5 Relative Density Tests Relative density tests were performed on selected, representative samples from the Mahomet Bedrock Valley deposit to determine the minimum and maximum densities. The results of these tests are presented in Table 2.5-34. These tests were performed according to ASTM D2049-69. 2.5.4.2.4 Chemical Tests Chemical analysis was performed on representative groundwater samples obtained from selected borings. The test results are presented in Subsection 2.4.13.1.3. The methods used were generally those given in "Standard Methods for the Examination of Water and Wastewater," Thirteenth Edition, 1971 (Reference 84). Calcium and magnesium were

determined by atomic absorption. 2.5.4.2.5 Chemical Analysis on Coal Samples Chemical analysis was performed on coal samples for the No. 8 and No. 7 coal beds in Boring P-38. The analysis was performed by the Illinois State Geological Survey and is presented in Table 2.5-3.

CPS/USAR CHAPTER 02 2.5-49 REV. 11, JANUARY 2005 2.5.4.2.6 Construction Stage Investigations This subsection presents static and dynamic engineering properties of the underlying materials encountered in the additional borings drilled and representative bulk samples taken as part of the engineering investigations and quality control program at the CPS site following the submittal of the PSAR. During the earthwork construction at the project site, there was a continuous program of monitoring the quality control of the earthwork construction and monitoring the soil conditions to confirm conformance with the design conditions. These monitoring activities involved boring programs, sampling programs, and testing programs. The data obtained from these programs provided documentation of the conformance of the work with the project earthwork specifications. Representative soil samples extracted from the test borings and stockpiles of borrow materials were subjected to laboratory tests to evaluate and document the physical characteristics of the in situ soils as well as the material placed as fill and backfill. The laboratory programs included the following tests: a. strength tests - 1. unconfined compression on disturbed and undisturbed samples; b. dynamic tests - 1. liquefaction tests; c. other physical tests - 1. Atterberg limit, 2. compaction, 3. in situ moisture and density determinations, 4. particle size analysis including hydrometer, and

5. relative density. The borings drilled after the writing of the PSAR are identified on the logs of borings as being logged by Sargent & Lundy. The results of the unconfined compression, in situ moisture and density, and Atterberg limit tests are presented to the left of the logs of borings. Other tests performed are indicated by symbols in the left-hand column of the logs of borings. The key to the test symbols is presented at the bottom of Figure 2.5-298. 2.5.4.2.6.1 Dynamic Tests 2.5.4.2.6.1.1 Laboratory Liquefaction Tests A series of laboratory liquefaction tests was performed on the Type B granular structural fill borrow material. Type B material is described in Subsection 2.5.4.5.1.5. Four bulk samples were taken from the material stockpile used in the construction operation in order to have CPS/USAR CHAPTER 02 2.5-50 REV. 11, JANUARY 2005 representative samples that closely correlate with the actual material used. The grain size characteristics are presented in Figure 2.5-348. To prepare the samples for testing, a preweighted amount of moist material with a water content of 7% was placed into a membrane lined forming mold 2.4 inches in diameter and 6 inches high in six equal layers. Each layer was compacted by a combination of vibratory and hand tamping methods to achieve the specified density. The densities of the samples ranged from 126.4 pcf to 127.2 pcf dry density. These values correspond to the average dry density at 85% relative

density. Carbon dioxide gas was then applied through the specimen for a period of time to aid in evacuating all of the air out of the specimen. Slight vacuum was then applied to maintain the shape of the specimen as the mold was removed and while micrometer measurements of the specimen diameter and height were obtained. The triaxial cell was assembled around the specimen, a small confining pressure applied, and the vacuum was released allowing water to slowly flow through the sample until it was nearly saturated. To fully saturate the specimen, back pressure was applied to the specimen through a water chamber with a confining pressure slightly higher than back pressure. The required effective confining pressure was then applied and testing was performed with values of Skempton's pore pressure parameter B exceeding 0.95. The liquefaction tests were performed under controlled stress conditions using a pneumatic sinusoidal wave loading system. After the specimen was fully consolidated, the drainage valve was closed and the specimen was subjected to a constant amplitude sinusoidal cyclic vertical load at a frequency of approximately 1 hertz. The traces of load, deformation and the pore water pressure with time were recorded on a strip chart recorder during testing. The criterion for liquefaction of the laboratory samples was defined as the number of cycles required to produce initial liquefaction. Initial liquefaction is defined as when the pore pressure first becomes equal to the effective confining pressure. For the Type B structural fill, the 5%

double amplitude strain and the initial liquefaction curves are shown in Figure 2.5-349. During cyclic testing, it was noted that the axial strain developed mostly at a small horizontal zone near the top portion of the specimen. It should be noted that both the recorded pore pressures and axial strains during cyclic loading cannot be attributed to uniform behavior of the entire specimen. Therefore, data points shown in Figure 2.5-349 for initial liquefaction and for 5%

double amplitude strain obtained during cyclic tests are conservative. A summary of the soil properties and the liquefaction test results is presented in Table 2.5-35. The results of the liquefaction tests are shown in the form of stress ratio (ratio of one-half of the cyclic vertical stress to confining pressure) versus the number of stress cycles required to produce liquefaction (or specified strain levels). 2.5.4.2.6.2 Other Physical Tests A laboratory testing program was established at the site to verify that the materials used during construction as structural fill and backfill met the specifications established for the materials. The following tests were performed on representative samples to maintain controls on the materials used and to ensure that the placed material conformed to the design considerations.

CPS/USAR CHAPTER 02 2.5-51 REV. 11, JANUARY 2005 2.5.4.2.6.2.1 Atterberg Limit Tests Representative soil samples were tested to evaluate their plasticity characteristics (ASTM D424-59 and ASTM D423-66). The results of these tests were used for classification and correlation purposes to verify the uniformity of the material. The Atterberg limit determinations are presented to the left of the boring logs. Table 2.5-37 includes a summary of the range of values for the Atterberg limit of the compacted cohesive material that was placed as fill material in the main plant and screen house areas. 2.5.4.2.6.2.2 Compaction Tests Compaction tests were performed on representative bulk samples of Wisconsinan till used as structural fill and backfill to maintain and verify the uniformity of the material. The compaction tests were performed in accordance with the American Society for Testing and Materials (ASTM) Test D1557. The results of the tests are presented in Figure 2.5-350 for material used as backfill in the plant and screen house areas. 2.5.4.2.6.2.3 In Situ Moisture and Density Determinations In situ moisture and density tests were performed on soil samples taken during drilling operations. The results of these moisture and density determinations are presented to the left of the boring logs. The moisture content determinations were performed according to ASTM D2216-66. In situ moisture (ASTM D-2216 and D-3017) and density (ASTM D-1556, D-2922, and D-2167) tests were also performed on the material placed as structural fill and backfill as part of the quality control program at the site. This is discussed in Subsection 2.5.4.5. 2.5.4.2.6.2.4 Particle Size Analyses Particle size distributions were determined for representative soil samples of both the fill materials and various soil strata to aid in classification and correlation of the physical soil properties as well as to verify the conformity of its gradation with the specifications. Figure 2.5-351 presents the statistical average grain-size distribution of the Type B materials used for structural fill and backfill in the plant area. These tests were performed according to ASTM

D422-63. 2.5.4.2.6.2.5 Relative Density Tests Relative density tests were performed on selected, representative samples of coarse-grained soils to determine the minimum and maximum densities for placement of the material as structural fill and backfill. The results of these tests are presented in Figure 2.5-377. These tests were performed according to ASTM D2049-69.

2.5.4.3 Exploration The subsurface soil, rock, and groundwater conditions at the station site (76 borings), at the ultimate heat sink (60 borings), at the dam site (60 borings), at the dam borrow area (17 borings), at the location of Section E-E' along the North Fork of Salt Creek (4 borings), and at the structural fill borrow areas (31 borings), were explored by drilling test borings at the locations indicated in Figures 2.5-14, 2.5-16, 2.5-271, 2.5-272, 2.5-273, 2.5-352, and 2.5-353.

CPS/USAR CHAPTER 02 2.5-52 REV. 11, JANUARY 2005 The borings taken in the PSAR stage were drilled with truck-mounted rotary wash or continuous-flight auger equipment by Raymond International, Inc., under the supervision of Dames & Moore. The rock was cored utilizing NX double-tube core barrels, which provide rock cores approximately 2 inches in diameter. Undisturbed soil samples suitable for laboratory testing were obtained using Dames & Moore Type U Sampler, a Pitcher sampler, and a 4-inch inside diameter double-tube core barrel (High Recovery Core Barrel). The Dames & Moore sampler is illustrated in Figure 2.5-354. This sampler is 3-1/4 inches in outside diameter and approximately 2-1/2 inches in inside diameter. The Pitcher sampler consists of a stationary thin inner barrel and a rotating outer barrel with a cutting bit, which is drilled into the soil. The stationary inner barrel has an outside diameter of 3.0 inches and an inside diameter of approximately 2.9 inches. The 4-inch inside diameter double-tube core barrel is very similar to the conventional double-tube rock core barrel. The core bit has an inside diameter of 4.0 inches and an outside diameter of 5.5 inches. Disturbed soil samples were extracted utilizing a standard split-spoon samplier approximately 2 inches in outside diameter. These samples were taken using the Standard Penetration Test procedure (ASTM D-1586). Additional boring programs were undertaken under the guidance of Sargent & Lundy personnel. The borings were drilled with truck-mounted rotary wash or continuous-flight auger equipment by Raymond International, Inc. Samples obtained during these programs were extracted utilizing a standard split-spoon sampler approximately 2 inches in outside diameter. These samples were taken using the Standard Penetration Test procedure (ASTM D-1586). Other sampling methods and equipment used included the 4-inch inside diameter double-tube core barrel, Pitcher sampler, Osterberg sampler, and Shelby tubes as described above. The borings taken for these programs are so noted on the logs themselves. A graphical representation of the soils and rock encountered in the borings, including standard penetration tests data, sampling, and coring information, is presented in Figures 2.5-19 through 2.5-270 and 2.5-439 through 2.5-450. The method utilized in classifying the soils and rock is described in Figure 2.5-355. The plans and profiles for the main plant excavations are discussed in Subsection 2.5.4.5. Attachment C2.5 presents the geologic mapping performed at the CPS site. A key to the sample symbols and the sampling information presented on the log of borings is shown in Figure 2.5-298. Piezometers were installed in the boreholes to observe ground-water conditions. They consist of 3/4-inch PVC pipe having an 18-inch-long porous stone at the bottom. A bentonite plug was installed above the porous stone which was enclosed in granular backfill. When the intended zone of percolation to be measured was large, 3/4-inch-to-2-inch-diameter perforated PVC pipe without a porous stone tip was used as the groundwater observation well. Nine additional piezometers were installed downstream of the main dam to monitor groundwater and seepage. They consisted of 1-1/2-inch diameter PVC pipe with the lower end plugged and the lower 5 feet slotted. A bentonite plug was installed above the slotted zone which was enclosed in granular backfill. A summary of the elevations at which piezometers were installed and observations of water levels are presented in Subsection 2.4.13.2.3. Falling-head-type field permeability tests were performed at the dam site and the CPS site. The results of these tests are presented in Table 2.5-38.

CPS/USAR CHAPTER 02 2.5-53 REV. 11, JANUARY 2005 2.5.4.4 Geophysical Surveys A program of integrated geophysical explorations was conducted at the station site, the dam site, and at Section E-E' along the North Fork of Salt Creek. This program consisted of the following field studies: a. a seismic refraction survey to evaluate the compressional wave velocities of bedrock and the materials overlying bedrock (results of this survey were also used to provide additional data to determine the depth to bedrock under the site); b. an uphole velocity survey to further define the compressional wave velocities of the materials overlying bedrock; c. a surface wave survey to determine surface wave types and characteristics; d. a downhole shear wave survey to evaluate the shear wave velocities of bedrock and of the materials overlying bedrock; and e. ambient noise studies to determine the predominant characteristics of ground motion due to background noise levels. The locations of these field studies are shown in Figures 2.5-356 through 2.5-358 and Figure 2.5-16. 2.5.4.4.1 Seismic Refraction Survey A seismic refraction survey was conducted by Dames & Moore at the station site along five seismic test lines for the total length of 6100 lineal feet. In addition, seismic refraction surveys were conducted at the dam site, and at Section E-E' along the North Fork of Salt Creek. The total length of these two additional surveys was 4000 lineal feet. The length of seismic test lines in the station site area was limited due to the proximity of two oil-product pipelines which pass through the area just south of the station site.

The seismic energy used in the survey was produced by explosive charges placed in drilled holes. These holes ranged in depth from 10 to 25 feet. The explosive charges ranged in size from 5 to 15 pounds of Nitramon-S (Du Pont).

The energy released by the explosives was picked up on vertically oriented geophones spaced at 100-foot intervals along the seismic test lines. The geophones, manufactured by Electro-Tech Labs, have a natural frequency of 14 cycles per second (Hertz) and are fitted with a spike to ensure proper coupling with the underlying soil. The energy impulse picked up by the geophones was recorded by an Electro-Tech Labs M4-E seismic amplifier coupled with an SDW-100 recording oscillograph. An Electro-Tech Labs ER-75-12A seismic recording amplifier was also used to record the energy impulses. The geophysical field crew consisted of two geophysicists, a licensed powderman with helper, and a driller with helper. The field work was performed from June 6 to June 23, 1972.

CPS/USAR CHAPTER 02 2.5-54 REV. 11, JANUARY 2005 The compressional wave velocities and the corresponding depths to the various subsurface layers under the site were evaluated by plotting the first arrival times of the seismic energy at each geophone against the distance of each geophone from the source of the seismic energy.

The time-distance data from each profile and the corresponding cross section of the subsurface layers for that profile are shown in Figures 2.5-359 through 2.5-365. 2.5.4.4.2 Uphole Velocity Survey An uphole velocity survey was performed by Dames & Moore at the CPS site, at the dam site, and at Section E-E' along the North Fork of Salt Creek. This provided a check on the compressional wave velocities measured during the seismic refraction surveys at each of these locations. Explosive charges of 1 pound of Nitramon-S were placed in drill holes to depths of 5 feet. These drill holes were offset from the boring to be studied by a distance of 25 feet. The seismic

energy released by the explosives was detected in the boring by twelve geophones affixed to a special cable (velocity cable). The energy was recorded on an Electro-Tech Labs ER-75-12A seismic recording amplifier. The results of the uphole velocity surveys are presented in Figures 2.5-366 through 2.5-368. The variation between the compressional wave velocities as determined by the uphole method and the refraction method is a result of the different techniques used in these methods, and the differences in the wave paths travelled by the seismic energy. The seismic refraction survey measures the compressional wave velocities over a lateral distance, whereas the uphole velocity survey measures the compressional wave velocities around an isolated point (the

boring). The uphole survey for Boring P-14 has not been grossly averaged. All points as shown are within a 2 msec deviation from the average velocity curve. A 2 msec deviation is considered normal, in that a 1 msec tolerance is allowed for the picking of the timebreak, and a 1 msec tolerance is allowed for the picking of the energy arrival. The interval in Boring P-14 from approximately 10 to 60 feet below the ground surface shows the largest deviation of the compressional wave velocities between these two methods (uphole and refraction). This is a result of the material within that layer. This material could show a slight increase in velocity with depth. However, because of the

+/- 2 msec error factor, this magnitude of velocity change in the uphole survey would not be accurately differentiated. The refraction work resulted in showing the higher velocity. A 100-foot geophone spacing is not too large to capture the effects of layering. The seismic refraction method is based on the compressional wave velocities and corresponding thicknesses of different layers. If a sufficient velocity contrast between two layers does not exist, if the layers are not sufficiently thick, or if velocities decline with depth, no refraction of a compressional wave will be noted. The uphole survey was performed on the site before the refraction survey. The interpretation of the uphole survey did not indicate any velocity inversions or sufficient velocity contrasts or layering thicknesses other than as shown. Therefore, the 100-foot geophone spacing was selected for the refraction work. Velocity measurements, other than those shown, cannot be determined for additional layers from the uphole survey. A common technique used in the petroleum exploration industry is to CPS/USAR CHAPTER 02 2.5-55 REV. 11, JANUARY 2005 integrate sections of an uphole velocity survey to obtain interval compressional wave velocities for different units. This technique is not possible, however, with the accuracy involved (

+/- 2 msec) for the uphole survey in Boring P-14. The geophysical program was established to determine the elevation of the bedrock beneath the site. The length of the refraction lines used for this purpose was not great enough to allow the delineation of any other major velocity units beneath the top of bedrock. The depth of the high velocity rock (velocity range from 9750 to 10,500 feet per second) in the station area is shown on the subsurface sections accompanying each time-distance plot. 2.5.4.4.3 Surface Wave Survey Surface wave surveys were conducted by Dames & Moore along a 2000-foot line in the CPS area, an 1800-foot line at the dam site, and a 1300-foot line at Section E-E' along the North Fork

of Salt Creek. The surface wave characteristics and shear wave velocities were computed from the recordings of two three-component geophones (Sprengnether Engineering Seismograph geophones) and eight one-component geophones. The Sprengnether geophones were placed 300 to 350 feet apart, and the one-component geophones were placed at 50-foot intervals between the Sprengnether geophones. The output of all the geophones was recorded by an Electro-Tech Labs M4-E seismic amplifier and an SDW-100 recording oscillograph.

The seismic energy for this survey was produced by small explosive charges detonated from 400 to 2000 feet from the geophone array. The characteristics of the surface waves for each site are listed in Tables 2.5-39 through 2.5-41. The characteristic frequency range is between 7 and 12 hertz. Significant amplification of seismic energy will probably occur only with this frequency range. It was also observed that the vertical component of particle motion attenuates much faster at the dam site than at the station site. This is probably a result of the different site conditions (subsurface) and the difference in the site geometry between the station site and the dam site. The surface wave survey provided data on shear wave velocities of the near-surface materials. These velocities confirmed the shear wave velocities measured by the downhole techniques. 2.5.4.4.4 Downhole Shear Wave Survey A downhole shear wave survey was performed by Dames & Moore at each site utilizing borings that had been drilled into bedrock. Two three-component low-frequency geophones (Mark Products L-1-3DS) were placed at different depths in each boring. Small explosive charges were detonated at varying distances from each boring. The resultant seismic energy was recorded by an Electro-Tech Labs M4-E seismic amplifier and an SDW-100 recording oscillograph. In addition, the seismic energy resulting from a technique of producing horizontally generated waves was also recorded. Recordings were made in each boring at successive 25-foot intervals.

CPS/USAR CHAPTER 02 2.5-56 REV. 11, JANUARY 2005 Shear wave arrivals are often masked by the compressional wave energy or the relatively large amplitude motion induced by surface and body wave systems. To overcome this difficulty, high-energy and low-energy recordings were made at each depth in the borings. The results of the downhole shear wave values and surface wave shear wave values are presented in Figures 2.5-369 through 2.5-371. These data are referenced to the subsurface conditions found at each site. 2.5.4.4.5 Ambient Vibration Measurements Measurements of the levels of ground motion due to background (ambient) vibrations were taken at each site at the locations shown in Figures 2.5-356 through 2.5-358. The measurements were taken when no equipment or drills were operating. A three-component VS-1200 Sprengnether Engineering Seismograph was used to record ambient ground motions. This seismograph has gain characteristics in the velocity mode of 20 inches/inch/second, the acceleration mode of 12 inches/inch/second/second, and the displacement mode of 200 inches/inch. A VS-1100 amplifier with a gain characteristic of 100 was used for all recordings. The resultant maximum gain level for the velocity mode is 2000; for the acceleration mode, 1200; and for the displacement mode, 20,000. The three components of ground motion measured were radial, vertical, and transverse. The seismometer was oriented facing north. The results of the ambient ground motion measurements are presented in Table 2.5-42.

2.5.4.4.6 Summary of Laboratory and Field Wave Velocity Measurements Compressional wave velocities derived from the laboratory dynamic triaxial tests and shockscope tests are summarized in Table 2.5-43.

Shear wave velocities derived from the laboratory resonant column tests are shown in Table 2.5-44. Compressional and shear wave velocities derived from the geophysical field data are summarized in Tables 2.5-45 and 2.5-46,

respectively. Comparing the laboratory and field wave velocity data shows that field and laboratory wave velocities are significantly different from each other. These differences occur because dynamic properties of soil and rock are strain-dependent. The laboratory dynamic triaxial tests and resonant column tests are usually performed with 5 to 10

-2 percent and 10

-2 to 10-4 percent of shear strain, respectively; whereas the field geophysical methods usually result in shear strains in the range of 10

-3 to 10-5 percent. Actual strong-motion earthquakes generally produce shear strains between 10

-1 to 10-3 percent. Therefore, the response calculations, wave velocities and dynamic properties derived from laboratory triaxial compression and resonant column tests at the shear strain of interest to the prevailing earthquake conditions are the most appropriate to be used. The shockscope tests were performed at zero confining pressure and at a level of strain considerably lower than most significant earthquakes. The wave velocities data as determined by the shockscope would therefore not be realistic and reliable for analysis.

CPS/USAR CHAPTER 02 2.5-57 REV. 11, JANUARY 2005 2.5.4.5 Excavations and Backfill 2.5.4.5.1 Station Site 2.5.4.5.1.1 Site Preparation Site preparation and earthwork for CPS Units 1 and 2 consisted of stripping, excavating, dewatering, and backfilling operations to attain a nominal station grade at approximately elevation 736 feet. Trees, brush, crops, grass, roots, and other deleterious materials were stripped from areas to be occupied by structures and from all areas that received fill. All topsoil was removed prior to general excavation operations. 2.5.4.5.1.2 Excavation The excavation for the main power station was an open excavation performed by heavy earth moving scrapers. The excavation was approximately 800 feet by 800 feet and it extended from existing grade to the Illinoian till of the unaltered Glasford Formation at approximately elevations 680 to 683. The maximum height of the construction slopes was approximately 56 feet in the excavation for the station. The limits of the excavation were extended a minimum of 20 feet beyond the limits of the structures themselves in order to ensure adequate foundation support within the zone of major stress concentrations. Based on the results of comprehensive engineering analyses, construction slopes for the major excavations were cut to a general slope of 1.5:1 (horizontal to vertical), but no steeper than 1:1 (horizontal to vertical). Figure 2.5-16 shows the location of Seismic Category I structures and exploration in the vicinity of these structures. Excavation details with geologic sections are shown in Figures 2.5-372 and 2.5-373. A plan of the station foundations is shown in Figure 2.5-374. 2.5.4.5.1.3 Dewatering The subsoil and groundwater conditions in the station area were such that no significant dewatering problems occurred. The rate of seepage into the excavations extending below the groundwater level was very low in the natural clayey till soils. However, more previous sand layers and seams did contribute to the rate of seepage. Dewatering was accomplished by a network of perforated metal pipe drains and ditches that collected the seepage at the periphery of the excavation. The seepage water was then drained from the excavation by gravity flow to the heat sink area to the west of the site. Water was not allowed to pond in the base of the plant excavation to permit base treatment and fill placement

under dry conditions.

CPS/USAR CHAPTER 02 2.5-58 REV. 11, JANUARY 2005 2.5.4.5.1.4 Excavation Base Treatment The base of the main plant excavation was on Illinoian till of the unaltered Glasford Formation.

Pockets of sand exposed in the base of the excavation were tested as described in the following. A comprehensive program of subgrade inspection and testing was initiated to verify the adequacy of the subgrade and compare the existing conditions with those postulated from the soil borings in the PSAR stage. Subgrade testing was performed as a part of the contractor's quality control program. Field testing on subgrade consisted of determining the in situ density using either the sand cone (ASTM D-1556) or nuclear densometer (ASTM D-2922) method. Laboratory testing consisted of the classification of soils by the unified classification method (ASTM D-2487) and determination of index properties using the applicable ASTM methods. These tests and the geologic mapping program described in Attachment C2.5 verified that the subgrade consisted of the Illinoian till of the unaltered Glasford Formation. The subgrade was considered adequate when the average in situ dry density exceeded a conservative value of 130 pcf for cohesive materials, or a minimum of 85% relative density (ASTM D-2049) was achieved at all test locations consisting of granular material. In areas where these requirements were not achieved, additional compaction and/or testing was performed to confirm the adequacy of the subgrade. Additional excavation was performed where subgrade conditions could not be improved. Some localized pockets of sand that could not be compacted to meet the requirements were removed and dental work was performed using a flyash backfill mixture. Figure 2.5-375 illustrates the areas that received the flyash backfill. Flyash backfill is a mixture of cement, flyash, sand, and water mixed in a central concrete batching plant. The backfill was transported and placed so as not to permit sedimentation. The material was tested in place and was considered to be acceptable when it had a strength which would yield a deflection of less than 0.25 inches under an applied load of 50 psi. A total of 90 in-place strength tests were performed on the bash placed as structural fill beneath the main power block. The minimum load used was approximately 55 psi and the maximum deflection recorded was 0.194 inches. Therefore, the bash placed beneath the power block is acceptable. The flyash mixture is described in Subsection 2.5.4.14. 2.5.4.5.1.5 Structural Fill and Backfill Fill placement commenced following subgrade approval. Controlled compacted granular fill, Type B material, was placed to bring the base elevation of the excavation up to the grade elevations for the foundations of the various plant structures. The dimensions and elevations of the foundation mats of the various structures, the type and thickness of the bearing strata, and the nature of the in situ soil underlying the bearing strata are summarized in Table 2.5-47. The details of the station structure foundations and the bearing strata are also shown in Figure 2.5-372. The soil exploration program for the Type B material involved the location of a borrow area from which the structural fill and backfill could be obtained. Aerial photographic interpretation, field CPS/USAR CHAPTER 02 2.5-59 REV. 11, JANUARY 2005 reconnaissance, soil exploration data, and published literature had been used in the borrow area survey. The original proposed borrow area shown in Figure 2.5-352 was investigated by the G-series borings illustrated in Figures 2.5-152 through 2.5-161. This area, discussed in the PSAR, was not used as a source of the structural fill and backfill for the main plant. Instead, another borrow area, shown in Figure 2.5-353, was considered to be more accessible and to have material more suitable for use as Type B structural fill. However, the initial testing performed on the samples taken from the proposed borrow area is still considered valid since the materials in the two borrow areas are of similar origin. The granular material that was used as Type B structural fill and backfill was obtained from the area designated in Figure 2.5-353. This area was investigated by drilling eleven auger Borings K-1 through K-8, K-11, K-12, and K-15, as shown in Figure 2.5-353. The logs of the borings are shown in Figures 2.5-243 through 2.5-253. The borrow area was located in alluvial deposits approximately 2.25 miles south of the station site. The fill was overlain by about 5 to 10 feet of clayey sand and clayey silt. The borrow material was a brown, poorly to well graded, fine to coarse sand with little silt, a trace of fine gravel, with coarse gravel and cobbles occasionally noted (SP, SW). A statistical average gradation of the Type B fill material is presented in Figure 2.5-351. After the overburden was stripped from the borrow area, a dragline operation was established to excavate and stockpile the material at the borrow area. Front end loaders were used to load the material into bottom dump and end dump trucks which hauled the material to the designated stockpile area adjacent to the main plant excavation. The intermediate stockpiling and handling aided in the mixing of the borrow material to produce a homogeneous mixture. Approximately one-half million cubic yards of this material were placed as structural fill and backfill in the plant excavation. Laboratory testing was performed on representative bulk samples of material used as Type B structural fill and backfill. The testing included: grain-size distribution, relative density, and dynamic triaxial tests. The testing is discussed in Subsection 2.5.4.2.6. Type B material used as structural fill to support foundation loads was placed in near-horizontal lifts. Each lift was compacted by a smooth wheel vibratory roller. The relative density was determined by the ASTM D-2049-69 test method. A statistical analysis of the in-place density tests for the Category I structural fill beneath the plant was performed to verify the compaction requirements. The average dry density and the average relative density of the fill were determined using the 4789 nuclear and sand cone tests performed on the granular fill. These averages were computed for each one-foot interval and for the total fill. The average relative densities for each one-foot interval range from 91.9% to 106.0%. The average dry densities varied from 129.5 lbs/ft 3 to 133.2 lbs/ft

3. The average relative density for the entire fill is 97.0%, with an average dry density of 131.5 lbs/ft

3. Figure 2.5-428 shows the distribution of the relative density for the in-place tests. Figure 2.5-429 shows the distribution of the dry density for the in-place tests. Figure 2.5-430 shows the distribution of the moisture content of the in-place tests.

CPS/USAR CHAPTER 02 2.5-60 REV. 11, JANUARY 2005 Of the 4789 tests, 175, or 3.7%, had relative densities less than the specified 85%. The locations of these 175 tests having lower densities are not concentrated at any one area or at any one elevation as shown in Figure 2.5-431. Figure 2.5-432 illustrates the grid pattern used to divide the power block fill area. A review of the areas surrounding these randomly scattered pockets of lower density material indicates that the preceeding and subsequent layers of fill were adequately compacted and met the specified density requirements. Also, the placement of additional fill above these areas of lower density increased the in-situ density in these areas.

Therefore, these lower density tests will not adversely affect the integrity of the fill. Based on this summary, the structural fill beneath the main plant structures at CPS is adequate to support the structural loads and meet the other criteria used in their design. A concrete mud mat was poured over the Type B structural fill to prevent rutting, erosion, and the provide a firm working area for the mat foundation. The specified compressive strength of the concrete used as the mud mat was 2,000 psi at 28 days. The actual compressive strength based on cylinder tests for all concrete used as a mud mat was significantly greater than the specified strength. Type B material used as backfill, at places other than to support foundations loads, was placed in near-horizontal lifts. Each lift was compacted by either a smooth wheel vibratory roller or vibrating tamping plates (for areas adjacent to structures). An analysis of the 3,479 in-place density tests taken on the granular backfill placed around the main plant was performed to summarize the data. Figure 2.5-482 shows the distribution of the dry density test results. The dry density ranged from 114.9 PCF to 147.2 PCF with an average value of 129.7 PCF. Figure 2.5-483 shows the distribution for the relative density test results. The relative density ranged from 56.1% to 134.8% with an average relative density of 94.5%. A total of 42 of the 3,479 tests performed did not meet the acceptance criteria for relative density. This represents 1.2% of these tests. Twenty-six of the 42 tests represent an area in the southwest corner of the main plant excavation. This area, approximately 100 feet by 300 feet, represents approximately 6 feet of fill, and has backfill material above and below it that met the criteria. A review of the tests in this area shows that the average relative density is 81.9%. Of the remaining 16 tests that did not meet the acceptance criteria, five tests had additional correlation tests performed that met the acceptance criteria and two tests had additional tests performed to average the values. For these two tests, the averages were acceptable. Based on this review the tests that did not meet the acceptance criteria represent only isolated areas that will not be detrimental to the integrity of the backfill. Type A cohesive material was used as backfill material over the granular material above approximately elevation 720 feet. Type A material was defined as having a plasticity index greater than 4.0 and no less than 45% passing the No. 200 sieve. The material was placed in near-horizontal lifts not exceeding 8 inches in loose thickness prior to compaction. An analysis of the 1,703 in-place density tests taken on the cohesive backfill placed around the main plant was performed to summarize the data. Figure 2.5-479 shows the distribution of the dry density test results. The dry density ranged from 112.9 PCF to 139.2 PCF with an average value of 124.0 PCF. Figure 2.5-480 shows the distribution of the moisture content. The moisture content ranged from 6.7% to 15.3% with an average value of 11.2%. Figure 2.5-481 shows the distribution of the percent compaction. The percent compaction ranged from 82.2%

to 104.0% with an average value of 93.0%.

CPS/USAR CHAPTER 02 2.5-61 REV. 13, JANUARY 2009 A total of 14 of the 1,703 tests performed on the cohesive material did not meet the acceptance criteria. This represents less than 1% of these tests. Only two tests did not meet the moisture requirement, and their values were only 0.1% outside of the acceptable limits. The remaining 13 tests did not meet the compaction requirement. Of these, only seven tests had a percent compaction less than 90%. These tests represent isolated areas and they will not be detrimental to the integrity of the backfill. Laboratory testing was performed on representative bulk samples of the material used as Type A backfill. The testing included: grain-size analysis, Atterberg limits, and Modified Proctor Compaction tests. This testing was discussed in Subsection 2.5.4.2.6. One bulk sample was

taken for every 6000 yd 3 of material placed. A summary of the properties on the in-place Type A material is presented in Table 2.5-37. The moisture-density relationships are presented on Figure 2.5-350. Flyash backfill was used as fill material in confined areas as shown in Figure 2.5-376 and Figure 4 of Question 241.8. The requirements for the fly ash backfill involve in-place testing with no compactive effort as previously stated in Subsection 2.5.4.5.1.4. The placement of all structural fill and backfill was monitored under a comprehensive quality control program which required in-place testing for each lift of material placed. The frequency of testing required for Type B material was four in-place tests for every 10,000 ft 2 of material per lift. The frequency of testing required for the Type A material was two in-place tests for every

10,000 ft 2 of material per lift. The construction of Unit 2 is cancelled. The majority of the excavation remains open, however, an administration building has been constructed at the north end of the excavation. Backfill behind the walls of Unit 1 has been extended and graded to drain water away from the open excavation. Erosion of the backfill due to incidental water will be prevented by the installation of a revetment composed of either gabions, cribbing, and/or a grout intrusion blanket similar to Fabriform. Water that collects within the excavation will be drained by gravity through the Unit 2 circulating water pipe to a juncture with a drain pipe that exits directly to the lake south of the

screen house (Q&R 241.18). Type B granular material from a borrow source (Borrow Area K) as shown on Figure 2.5-353 by the K-series borings, was used as structural fill and backfill for plant structures. In the PSAR the proposed source for the structural fill material was a borrow source (Borrow Area G) as shown on Figure 2.5-352 by the G-series borings. A composite sample from Borings G-18, G-19, and G-20 was used to determine static and dynamic strength properties for the structural fill material. Because of the similar nature and geologic source of these two borrow areas, the strength properties for these tests were also used for the material from Borrow Area K. These properties are: Consolidated-Undrained Triaxial Test Data - USAR Table 2.5-15 Dynamic Triaxial Compression Test Data - USAR Table 2.5-24

Resonant Column Test Data - USAR Table 2.5-28 Dynamic Triaxial Compression Tests - USAR Figures 2.5-302 to 2.5-311 CPS/USAR CHAPTER 02 2.5-62 REV. 13, JANUARY 2009 Parameters for Analysis of Rock-Soil-Structure Interaction - USAR Table 2.5-48 (for compacted structural fill) Consolidation Test - USAR Figure 2.5-338. Both the material used for structural fill from Borrow Area K and the material tested from proposed Borrow Area G were from alluvial deposits in the Salt Creek Valley. Revised Figure 2.5-14 shows the location of the two Borrow Areas G and K. The grain size curves for the structural fill material used and the material used for the strength property tests are shown on revised Figure 2.5-351. Both of these curves show the material as a well graded (Cu = D 60/D10 > 4) sand with the material used as fill being slightly coarser than the material tested. The use of test specimen from the proposed borrow area for static strength and consolidation properties should be conservative because the coarser material of the fill should have a greater angle of internal friction at corresponding relative densities. The dynamic properties should be similar because the grain size distributions are relatively close and the pore pressure response due to dynamic loading will be similar. The densities of the material were also very close. For the test material from Borrow Area G, a relative density equal to 91.0% corresponds to a dry density of 129.4 pcf where the average dry density of 129.1 pcf for the structural fill corresponds to an average relative density equals 91.0% based on the average minimum and maximum densities of the fill material (Q&R 241.1). 2.5.4.5.1.6 General Fill Beyond the 40-foot limit of the Category I backfill in the plant excavation, granular material was placed. This material was classified as Type B1 with the same soil characteristics as Type B material except that the gradation for the No. 200 sieve allowed for a maximum of 13% passing instead of only 10%. The placement of the general fill was similar to that of the Type B backfill except that the in-place density testing requirement was relaxed to a frequency of one test per

10,000 ft 2 per lift. Type C cohesive backfill material was also placed beyond the 40-foot limit of the Category I backfill. This backfill was a continuation of the lifts of the Type A material placed as Category I backfill described in Subsection 2.5.4.5.1.5. A description of the Type C material is presented in Subsection 2.5.4.5.2.5. The in-place density test requirement was relaxed to one test per

10,000 ft 2 per lift. The analysis of the in-place density test results for both the granular and cohesive general fill have been included in the respective backfill summaries in Subsection

2.5.4.5.1.5. Alternative IDOT CA-6 type backfill material may be used in non-safety-related fill applications associated with the RAT and ERAT transformer. 2.5.4.5.2 Circulating Water Screen House 2.5.4.5.2.1 Site Preparation Site preparation and earthwork for the circulating water screen house consisted of the same operations as described in Subsection 2.5.4.5.1.1.

CPS/USAR CHAPTER 02 2.5-63 REV. 11, JANUARY 2005 2.5.4.5.2.2 Excavation Figures 2.5-378 and 2.5-379 show the details of the foundation excavation for the construction of the circulating water screen house. The bottom of the foundation mat of the circulating water

screen house was approximately at elevation 653 feet. At the screen house location, the top of the Illinoian till was at approximately elevation 660 feet. The base of the excavation was established in sound Illinoian till of the unaltered Glasford Formation. The depth of excavation was approximately 65 feet and 35 feet on the east and west sides, respectively. Construction slopes of the 3:1 and 1:1 (horizontal to vertical) for the excavations in the Salt Creek alluvium of the Henry Formation and Illinoian glacial till of the unaltered Glasford Formation, respectively, were adequate to ensure the stability of the slopes. Construction slopes of 2.0-1.5:1 (horizontal to vertical) in the Wisconsinan till of the Wedron Formation and interglacial zone materials of the weathered Glasford Formation were adequate to ensure the stability of the slopes in these materials. 2.5.4.5.2.3 Dewatering Dewatering of the screen house was accomplished by providing a series of open drainage ditches at the outer limits of the excavation that collected the seepage water at sumps located in two corners of the excavation. The water was then pumped out as it became necessary to allow earthwork operations to be performed under dry conditions. The ditches were later backfilled with cohesive material. 2.5.4.5.2.4 Excavation Base Treatment Pockets of loose sand exposed in the bottom of the excavation were removed to establish the foundation on sound Illinoian till of the unaltered Glasford Formation. Subgrade verification was performed as part of a quality control program similar to that described in Subsection 2.5.4.5.1.4. Subsequent to subgrade testing, a concrete mud mat, 6 inches in thickness, was placed over the Illinoian till to protect the subgrade and to provide a dry working surface for the foundation

work. 2.5.4.5.2.5 Structural Backfill Material placed as structural backfill around the screen house consisted of cohesive material (Type C) excavated from the main station construction area and heat sink borrow area. Type C material was defined as having a plasticity index greater than 4.0 and no less than 45% passing the No. 200 sieve. It was placed and compacted to a minimum of 90% of the maximum dry density as determined by the ASTM D-1557 test method. The material was placed in near horizontal lifts prior to compaction. Laboratory testing was performed on representative samples used for the earthwork operations in the heat sink as well as on the material used specifically as backfill around the screen house. The moisture-density relationship for the material is shown in Figure 2.5-350. A summary of the properties of the in-place materials is the same as that presented in Table 2.5-37.

CPS/USAR CHAPTER 02 2.5-64 REV. 11, JANUARY 2005 Flyash backfill was used for dental backfill work around piping and to backfill the sumps as shown in Figure 2.5-376. A comprehensive quality control program was also followed for the backfill operations with two in-place density tests per 10,000 ft 2 per lift being required. An analysis of the 1,742 in-place density tests taken for the cohesive backfill was performed to summarize the data. Figure 2.5-476 shows the distribution of the dry density test results. The

dry density ranged from 116.3 PCF to 140.6 PCF with an average value of 125.0 PCF. Figure 2.5-477 shows the distribution of the moisture content test results. The moisture content ranged

from 5.6% to 15.2% with an average value of 11.0%. The percent compaction ranged from 87.4% to 105.6% with an average percent compaction of 93.7%. Figure 2.5-478 shows the distribution of the percent compaction. Of these 1,742 tests, 98 did not meet the acceptance criteria for either the percent compaction or moisture content. This represents 5.6% of the screen house backfill tests. However, 69 of the 98 tests represent the area where the SSWS pipeline enters the screenhouse. In this area, a 95% degree of compaction was required. The tests performed in this area were evaluated and the average percent compaction for this area was greater than 93% and was considered to be acceptable. The reduction in the compaction effort will not be detrimental to the function of this fill. Of the remaining 29 of the 98 tests, only four had a percent compaction less than 89%. One of these tests had two other tests performed near the same location with the average percent compaction of the three tests being greater than 90%. Six of these 29 tests did not meet the moisture content acceptance criteria only. However, their moisture contents were generally within 1.5% of the acceptable range, and will not be a detriment to the fill. In conclusion, the 98 tests that did not meet the acceptance criteria were found to represent only small areas which will not be detrimental to the integrity of the backfill. 2.5.4.5.3 SSWS Outlet Structure and Pipelines 2.5.4.5.3.1 Site Preparation Site preparation and earthwork for the shutdown service water system (SSWS) outlet structure and pipelines consisted of the same operations as described in Subsection 2.5.4.5.1.1. 2.5.4.5.3.2 Excavation The excavation for the SSWS outlet structure extended from the existing grade to the Illinoian till of the unaltered Glasford Formation approximately at elevation 655 feet. The excavated slopes from elevation 655 feet to 662 feet were near vertical and were approximately 5 feet from the structure itself. The slopes above elevation 662 feet were cut back on a 2:1 (horizontal to vertical) for construction purposes. The final slope configuration around the SSWS outlet structure is discussed in Subsection 2.5.5.1.2. The excavation and structural fill placed beneath the structure is illustrated in Figure 2.5-381. Excavation was performed along the SSWS pipeline alignments between the screen house and the station site and between the outlet structure and the station site. A longitudinal subsoil profile along the SSWS pipeline is presented in Figures 2.5-486 and 2.5-487. Typical CPS/USAR CHAPTER 02 2.5-65 REV. 11, JANUARY 2005 transverse sections illustrating the concrete mudmat, flyash mixture, pipe, and backfill materials are shown on Figure 2.5-488. Zones of soft and loose material were removed as indicated by overexcavation beneath the pipeline as shown on Figures 2.5-486 and 2.5-487.

(Overexcavation is considered to be any excavation greater than 1.5 feet below the bottom of the lower pipe.) Minor seepage into the pipeline excavation was pumped as it became necessary. This excavation was normally dry after rain. 2.5.4.5.3.3 Dewatering Minor seepage into the excavation was diverted around the outer limits of the outlet structure excavation by open ditches. The water was drained by gravity away from the excavation into a larger collector ditch from which the water was pumped as necessary. 2.5.4.5.3.4 Excavation Base Treatment The base of the excavation for the SSWS outlet structure was established on sound Illinoian till. Pockets of loose material were removed prior to subgrade testing and approval. A concrete mud mat, with a minimum thickness of 4 inches, was placed on the approved subgrade for the outlet structure to protect it from exposure. Soft material encountered immediately behind the outlet structure was overexcavated and Type A cohesive material was placed there. A total of 5 feet of Type A material was placed immediately behind the structure to replace the exca vated soft materials. An analysis of the 50 in-place density tests performed on the cohesive backfill around the outlet structure was made to summarize the data. The dry density of the material ranged from 122.9 PCF to 131.0 PCF with an average dry density of 126.5 PCF. Figure 2.5-468 illustrates the distribution of these values. The moisture content of the material ranged from 7.5% to 13.3% with an average of 11.0%. Figure 2.5-469 shows the distribution of the moisture content. The percent compaction ranged from 91.9% to 98.3% with an average value of 94.9%. Figure 2.5-470 shows the distribution of the percent compaction for these tests. Type A material is described in

Subsection 2.5.4.5.1.5. A concrete mudmat having a minimum thickness of 4 inches was placed beneath the SSWS pipeline either over the approved subgrade or structural fill along the pipeline. 2.5.4.5.3.5 Structural Fill and Backfill Type B granular fill material, as discussed in Subsection 2.5.4.5.1.5, was placed as structural fill directly over the mudmat beneath the outlet structure from approximately elevation 655 feet to 662 feet. This material was placed in near horizontal lifts. An analysis was performed on the 15 in-place density tests performed on the Type B granular fill. The dry density of this material

ranged from 127.4 PCF to 134.2 PCF with an average dry density of 131.1 PCF. Figure 2.5-466 shows the distribution of the dry density test results. The relative density, as determined by ASTM D-2049, ranged from 89.8% to 99.6% with an average relative density of 94.9%. Figure 2.5-467 shows the distribution of the relative density test results. All of these tests met the acceptance criteria of a minimum of 85% relative density. A thin concrete seal was placed over the Type B material to protect it from runoff water. Between the elevations of 662 feet and 669 feet, fly ash mixture backfill was placed and tested as described in Subsection 2.5.4.5.1.4. A 12-inch thick apron of the fly ash mixture backfill was CPS/USAR CHAPTER 02 2.5-66 REV. 11, JANUARY 2005 also placed along the two side walls of the outlet structure. Four in-place strength tests were performed on the fly ash mixture beneath the SSWS outlet structure. The maximum deflection was 0.022 inches for a load of 63.6 psi. This is less than the allowable deflection of 0.25 inches for a 50 psi load. A total of 24 in-place tests were performed on the fly ash mixture placed along the SSWS pipeline. A load of 71.7 psi was used for all of these tests with a maximum deflection of 0.174 inches being recorded. Therefore, the tests performed for the SSWS pipeline and outlet structure are acceptable. Flyash mixture backfill was placed around the SSWS piping as shown on Figure 2.5-488. Structural backfill was then placed and compacted over the pipes. Type B granular material was used as fill around the lower pipes immediately adjacent to the main plant structures. A summary of the 59 in-place tests performed in this area was made to summarize the data. The dry density of this fill ranged from 121.6 PCF to 132.7 PCF with an average value of 126.6 PCF. Figure 2.5-471 shows the distribution of the dry density test results. Figure 2.5-472 shows the distribution of the relative density test results. The relative density ranged from 85.6% to 118.0% with an average value of 100.8%. All of these tests met the acceptance criteria of 85% relative density. Cohesive material was used as fill around the SSWS pipeline in all the remaining areas. An analysis of the 524 in-place tests taken on the cohesive material was performed to summarize the data. Figure 2.5-473 shows the distribution of the dry density test results. The dry density

ranged from 116.2 PCF to 133.8 PCF with an average dry density of 122.3 PCF. Figure 2.5-474 shows the distribution of the moisture content for the tests. The moisture content ranged from 6.2% to 14.2% with an average value of 11.1%. Figure 2.5-475 shows the distribution for the percent compaction. The percent compaction ranged from 89.1% to 101.0% with an average value of 94.4%. Only seven of the 524 in-place density tests did not meet the acceptance criteria for percent compaction of this fill material. One of these tests also did not meet the moisture acceptance criteria. These failing tests represent 1.3% of the tests performed for the pipeline. As previously stated, the lowest percent compaction recorded was 89.1%. Also, these seven failing tests represent only isolated areas along with pipeline.

Therefore, the material represented by these tests will not be detrimental to the integrity of the pipeline fill. Section C-C on Figure 2.5-488 illustrates the use of the flyash mixture as it was placed within 15 feet of the bends in the SSWS pipeline. The flyash mixture was used as bedding and placed vertically up to 1/6 of the diameter of the pipe. Styrofoam, 6 inches in thickness, was placed between the flyash mixture bedding to made the bedding for each pipe independent of each other. Structural cohesive fill and backfill was then placed and compacted as previously

discussed. Information follows for the buried shutdown service water system (SSWS) piping outdoors. There is no ECCS piping buried outdoors. (Q&R 241.8)

(a) A longitudinal subsoil profile along the SSWS pipeline is presented in Figures 2.5-486 and 2.5-487. The excavation line for the pipeline is illustrated. The zones of soft and loose material that were removed are shown as over-excavation beneath the pipeline.

(Overexcavation is considered to be any excavation greater than 1.5 feet below the bottom of the lower pipe.)

CPS/USAR CHAPTER 02 2.5-67 REV. 11, JANUARY 2005 (b) Transverse cross sections showing the pipe, concrete mudmat, and all backfill materials are presented in Figure 2.5-488. (c) The details of the backfill placement near the connection between pipes and structures are shown in Section F-F of Figure 2.5-488. The estimated total settlement for the structure (Diesel Generator Building) where the SSWS pipes enter is 1 inch. This settlement is based on the estimated total settlement of the structure (Figure 2.5-436) beginning January 1, 1979. This is the approximate date of the connection of the first pipe to the structure. If it is assumed that the pipeline will not settle, the estimated differential settlement between the pipeline and structure will also be approximately 1 inch. (d) A figure showing the fly ash placement from the SSWS outlet structure and along the pipeline to the main plant is shown in Figure 2.5-496. (Q&R 241.8). 2.5.4.6 Groundwater Conditions A discussion of the history of the groundwater conditions, monitoring of piezometers, and groundwater conditions used in analyses is presented in Subsection 2.4.13. A discussion of the control of groundwater and seepage in the open excavations is presented in Subsection 2.5.4.5.1.3, 2.5.4.5.2.3, and 2.5.4.5.3.3 for the main plant, screen house, and outlet structure, respectively. 2.5.4.7 Response of Soil and Rock to Dynamic Loading The parameters utilized on soil-rock-structure interaction analyses are presented in Table 2.5-48. The static soil properties presented in this table were based on evaluation on laboratory consolidation and triaxial test data. The strain dependent dynamic moduli and damping values were evaluated on the basis of geophysical results and laboratory dynamic triaxial and resonant column tests. The selected design parameters reflect both the results of the tests performed during the PSAR investigation and properties previously developed for similar soils. The present design of walls and slabs enveloped the conditions of Unit 2 being presented and Unit 2 structures being deferred. In developing the soil-spring constants for soil-structure interaction being performed in response to Question 220.26, the overburden and confining pressure under the Unit 1 mat have been used. However, under the present condition where only one unit is constructed, the confining pressure under the area of Unit 2 excavation is smaller. This smaller confining pressure will lead to lower shear modulus values for the 20-foot-thick structural fill layer. A sensitivity study performed on soil properties in answer to Question 220.15 shows that lower modulus values lead to lower structural responses (Q&R 241.19). 2.5.4.8 Liquefaction Potential 2.5.4.8.1 Structural Fill The liquefaction potential of the structural fill beneath the structure was evaluated on the basis of the simplified procedure described by Seed and Idriss (Reference 85). The procedure is based on both theoretical considerations and descriptions of site conditions where liquefaction CPS/USAR CHAPTER 02 2.5-68 REV. 11, JANUARY 2005 was known to have occurred or not to have occurred under earthquakes of known or estimated magnitudes. The liquefaction potential of a granular soil deposit is related to: a. the grain-size characteristics of the sands, b. the relative density, c. the position of the groundwater table, d. the intensity and duration of ground shaking, and

e. the number of significant stress cycles produced by the earthquake. Results of the evaluation are presented in Table 2.5-49 for various depths in the structural fill beneath the structure. Evaluation is based on the safe shutdown earthquake postulated for the site (Subsection 2.5.2.6). The number of significant stress cycles is anticipated to be five cycles according to a correlation between the significant stress cycles and magnitude of earthquake established by Seed et al. (Reference 85). However, in the analysis ten significant cycles were conservatively used for the safe shutdown earthquake. The structural granular fill Type B material was compacted to a minimum relative density of at least 85% determined by ASTM D-2049 method of compaction. The maximum shear stresses were computed assuming that the soil behaves as a rigid body. The rigid body stresses were corrected using a stress reduction coefficient to account for the fact that the soil actually behaves as a deformable body. The average equivalent uniform shear stress during the earthquake is estimated to be 65% of the computed maximum shear stress.

The stresses required to produce liquefaction in ten cycles were computed using the stress ratio (ratio of one-half of the cyclic vertical stress to confining pressure) obtained from laboratory liquefaction tests. The laboratory liquefaction test results are shown in Figure 2.5-349. Because of the difference between the field and laboratory stress conditions and the limitations in testing equipment and procedures, a correction factor of 0.70 was applied to the laboratory values to obtain the cyclic shear stresses required to produce liquefaction in the field (Reference

86). The liquefaction potential of the structural fill was evaluated by computing the factor of safety with respect to liquefaction at various depths during the safe shutdown earthquake. The factor of safety is defined as the ratio of the cyclic shear stress required to produce liquefaction to the average uniform cyclic shear stress induced by the earthquake. The calculations are based on ten significant stress cycles. The water table was assumed to be located at elevation 730 feet, which is 6 feet below grade elevation. The results of the evaluation shown in Table 2.5-49 indicate that the factor of safety with respect to initial liquefaction during the safe shutdown earthquake is 2.03 for the structural fill. Therefore, there is no liquefaction potential for the structural fill beneath the structure which is to be compacted to a minimum relative density of least 85% as obtained by ASTM D-2049 method of compaction. Question 241.9 concerning the liquefaction potential of the natural material in the vicinity of ECCS piping is assumed to apply to buried shutdown service water system (SSWS) piping outdoors. There is no ECCS piping buried outdoors.

CPS/USAR CHAPTER 02 2.5-69 REV. 11, JANUARY 2005 The material used as structural backfill around the SSWS piping consisted of concrete, bash, and Type A cohesive material. These materials are not susceptible to liquefaction and thus not likely to liquefy. As discussed in Attachment C2.5, Geologic Mapping, the subgrade for the SSWS piping consisted of the Wisconsinan Till of the Wedron Formation. This till consists of cohesive material with isolated and discontinuous pockets of sand and silt randomly distributed within the till. The cohesive material is not susceptible to liquefaction and thus not likely to liquefy. Some small isolated sand pockets were encountered during the excavation for the SSWS pipeline. These are shown on Figure C2.5-23. In-place density tests were performed in these areas. The results of these tests indicate that these sand pockets have an in-place relative density greater than 82.5%. Based on the facts that the sand pockets are confined, discontinuous, and that they have an in-situ relative density greater than 82.5%, they are not considered to be very likely to liquefy (Q&R 241.9).

The minimum factors of safety for sand elements were obtained by comparing the shear stresses required to cause single-amplitude shear strain of 5% (see Figure 2.5-413) with the equivalent shear stresses induced by the earthquake. It can be seen from Table 2.5-68 that the 5% strain occurs before the initial liquefaction starts. Therefore, the criterion of 5% strain is conservative, and it assures that the sand elements will not liquefy during the earthquake (Q&R

241.16). 2.5.4.8.1.1 Structural Fill Subjected to New Madrid Type Earthquake The liquefaction potential of the structural fill was analyzed in the same way for the New Madrid type earthquake (Subsection 2.5.2.6) as for the safe shutdown earthquake. The details of the analyses are the same in both cases except for the earthquake parameters. The changes in the parameters for the analyses of the New Madrid type earthquake are: a. a maximum ground acceleration 0.13g, and b. the number of significant stress cycles was assumed to be 30. The results of the evaluation shown in Table 2.5-50 indicate that the minimum factor of safety with respect to initial liquefaction during the New Madrid type earthquake is 2.14 for the structural fill. 2.5.4.8.2 Sand Lenses The soils under the station site above elevation 683 feet were removed except in a few isolated areas where sound Illinoian till of the unaltered Glasford Formation was encountered at approximately elevation 686. The excavation was then filled with Type B granular material compacted to a minimum relative density of 85% as determined by the ASTM D-2049 test method and described in Subsection 2.5.4.5.1.5. The sand lenses below elevation 680 feet were examined for liquefaction potential by evaluating the soil borings under the station mat including the seven additional soil borings, P-33A and P-50 through P-55, drilled specifically for this purpose. Further discussion of the borings within the station mat is given in Subsections

B2.5.2 and B2.5.2.1.

CPS/USAR CHAPTER 02 2.5-70 REV. 11, JANUARY 2005 Based on the physical data indicated on the boring logs, which includes standard penetration test values, relative density, grain size analyses, confining pressures, material type, and in situ dry density, it is concluded that the sand lenses are not susceptible to liquefaction. Therefore, no settlements due to liquefaction are anticipated. The facts leading to this conclusion are discussed in detail in Attachment B2.5. 2.5.4.9 Earthquake Design Basis The response spectra defined in Subsections 2.5.2.6 and 2.5.2.7, and presented in Figures 2.5-296 and 2.5-297, specify the earthquakes which are used in the design and analysis. 2.5.4.10 Static Stability 2.5.4.10.1 Foundation Exploration and Testing Program Geologic, seismologic, and foundation conditions of the site were evaluated by the boring program presented in Subsection 2.5.4.3. Subsection 2.5.4.2 describes in detail the static and dynamic engineering properties of the materials underlying the site. Figure 2.5-271 shows the location of the borings on the site. Borings used in the analysis of station foundation conditions are listed in the following tabulation: BUILDING NO. OF BORINGS BORING NUMBER Fuel, containment, and auxiliary buildings (Unit 1) 5 P-14, P-37, P-38 P-54, P-55 Fuel containment, and auxiliary buildings (Unit 2) 3 P-18, P-29, P-30 Turbine building (Unit 1) 5 P-39, P-40, P-41, P-52, P-53 Turbine building (Unit 2) 7 P-15, P-31, P-32, P-33, P-33A, P-50,

P-51 Radwaste 2 P-10, P-36 Control 1 P-35 Immediately adjacent to station area 7 P-7, P-9, P-11, P-34, P-42, P-43,

P-47 Engineering analysis has shown that the supporting capacity of the Illinoian till and underlying lacustrine and pre-Illinoian soils is far in excess of the pressures imposed by the mat foundations under static and dynamic loading conditions. Settlement analysis indicates that the anticipated foundation loads imposed on the Wisconsinan till and the interglacial zone materials which overlie the Illinoian glacial till would cause excessive settlement of the station complex. These soils were removed and replaced with a controlled, compacted, granular Type B fill as discussed in Subsection 2.5.4.5. The compacted CPS/USAR CHAPTER 02 2.5-71 REV. 11, JANUARY 2005 fill bears directly on the underlying hard Illinoian glacial till strata. The hard Illinoian glacial till consists of gray to brown clayey silt with occasional gravel-sized particles. The underlying lacustrine and pre-lllinoian soils are hard and consist of clayey and sandy silts with occasional gravel-sized particles. The underlying lacustrine and pre-Illinoian soils also exhibit high density and high shear strength characteristics. The approximate thickness of the Illinoian till is 120 feet and the combined thickness of the underlying lacustrine and pre-Illinoian soils is approximately 60 feet. These soils are immediately underlain by thinly bedded Pennsylvanian limestone and shale. 2.5.4.10.2 Bearing Capacities The fill beneath the station foundations was constructed using selected granular materials (see Subsection 2.5.4.5.1.5) placed using controlled compaction procedures. The compacted fill has excellent supporting capacity to transmit the anticipated structural loads to the underlying hard Illinoian glacial till without excessive foundation settlements. Ultimate bearing capacities and factors of safety for the station mat foundations, the circulating water screen house, and the ultimate heat sink outlet structure are presented in Table 2.5-63.

These factors of safety were calculated by conventional bearing capacity analyses assuming a local shear failure condition. It was assumed that the subsoil beneath the foundations is uniform and the mats under the various components of the station are structurally independent with respect to foundation loading and support. The total static foundation loads refer to total dead loads and equipment operating loads. Since the rigidity of the soil will increase during dynamic loading which is usually associated with smaller strain, the ultimate bearing capacities of the underlying cohesive soils will be larger under dynamic loading than for static conditions; and as a result, the factors of safety under seismic loading conditions would not be substantially reduced. The interaction of the subgrade and the mat foundations was investigated to evaluate the effects of slab deflection under static loading conditions. A static modulus of subgrade reaction of 25 to 300 psi was utilized in the analysis of the mat foundations. The CPS power block complex is supported on a monolithic basemat. For combined SSE and pool dynamic loads (SRV and LOCA), the maximum foundation-bearing pressure and displacement under the mat can be found in calculation SDQ12-21DG06. 2.5.4.10.3 Settlement The results of consolidation tests on soil samples obtained from various elevations at the station site indicate that the Wisconsinan glacial till and the interglacial zone materials were more compressible than the underlying Illinoian glacial till. Consequently, mat foundations established in the near-surface solid would undergo excessive settlement when subject to moderately high foundation pressures. These materials were, therefore, considered unsuitable as bearing strata for the mat foundations, and were removed by excavating to the top of the Illinoian glacial till. Type B controlled compacted granular fill was placed over the subgrade to bring the bottom of the excavation up to foundation grade. The results of consolidation test data are presented in Figures 2.5-320 through 2.5-346 and summarized in Table 2.5-62. Consolidation tests were performed on representative samples as described in Subsection 2.5.4.2.3.2. The test results indicate that the Illinoian and pre-Illinoian CPS/USAR CHAPTER 02 2.5-72 REV. 11, JANUARY 2005 tills are overconsolidated. Estimates for the maximum past consolidation stress based on Casagrande method of construction (Reference 87) indicate that the preconsolidation pressure of the Illinoian glacial till lies between 14.0 and 25.0 ksf (Table 2.5-62). Also, vertical stresses in the subsoil due to structural loadings do not exceed the difference between the maximum past preconsolidation pressure and the existing effective overburden pressure. Hence, values of the recompression index, C r, were used to estimate the settlement. The soils were divided into eight strata in the settlement analysis. The recompression indices assigned to each soil stratum are presented in Table 2.5-66. For computing the rate of settlement, an average value of coefficient of consolidation of 11.4 sq. ft. per day is used for the entire soil profile. Settlement analyses have been performed to compute the settlement of plant structures according to the construction sequence. Excavation for the plant structures began in the fall of 1975 and was completed in about nine months. Backfill of the controlled compacted fill and construction of Unit 1 then followed. For analysis purposes, a three-month period for the placement of controlled compacted fill and a four-year period for the construction of Unit 1 structures are used. It is further approximated that the construction period of the mat foundation of Unit 1 is one year. The construction of Unit 2 has been cancelled. The final settlement of Unit 1 was less than that shown in Figure 2.5-433. The grade elevation of the station site is approximately 736 feet. The design groundwater level is assumed to be elevation 730 feet (Subsection 2.4.13). The foundation elevations and static loads for each of the plant structures are shown in Table 2.5-63. The construction time sequences are divided into excavation, placement of the controlled compacted fill, and construction of plant structures. There is no settlement due to the dewatering process employed during construction (Section 2.5.4.5.1.3). The settlement due to applied building loads is computed based on the gross foundation loads minus the uplift

pressures. The SETTLE computer program described in Appendix C is used to compute settlements. SETTLE uses the comparison index method to perform the settlement analysis. The stress increment due to applied external loads is computed using Bouissinesq's formula. The settlement due to the applied building loads is computed by assuming that the structural foundation system is either completely rigid or completely flexible. The settlement is taken as the average of the results obtained from these two cases. For the completely flexible case, the rigidity of the foundation and superstructure system is neglected. The effective foundation pressure is applied at the foundation level directly on top of the soil. For the completely rigid foundation case, the distribution of contact pressure due to foundation rigidity is taken into account by considering linear settlement. An iterative procedure is used to make the settlement pattern of the foundation and the subsoil compatible. This iterative procedure is included in the SETTLE program. The settlement time histories for the main plant are computed using the one-dimensional consolidation theory. Each loading (i.e., excavation, structural fill, applied building loads) is applied independently and the resulting settlement time history is obtained by superposition. The initiation of settlement of the plant structures is taken at the completion date of the construction of the Unit 1 mat foundation. This point on the computed settlement time history curve, therefore, marks the origin of the datum line for defining plant settlement. The computed CPS/USAR CHAPTER 02 2.5-73 REV. 11, JANUARY 2005 time histories of plant settlement at four settlement points are shown in Figures 2.5-434 through 2.5-437. The contours of the computed final settlement for the main plant area are shown in Figure 2.5-433. The differential settlements between adjacent buildings are not presented because the entire power block is on one monolithic mat foundation and various buildings are also rigidly connected with continuous shear walls and diaphrams. An instrumentation system of settlement monuments was established at various locations throughout the main plant buildings to monitor the settlement. This system is discussed in Subsection 2.5.4.13. The locations of those settlement monuments are shown in Figure 2.5-382. The graphical plots of measurement readings recorded to date (January 1984) for those settlement monuments are presented in Figure 2.5-438. A comparison between the calculated final settlement and the measured settlements to date (January 1984) for all the settlement monuments is presented in Table 2.5-67. The calculated and measured settlement time history values are compared for settlement monuments C4 (Containment Building, Unit 1), T2 (Turbine Building, Unit 1), D3 (Diesel Generator Building), and R3 (Radwaste and Off-Gas Buildings), as shown in Figures 2.5-434 through 2.5-437 respectively. It should be noted that the first measurement readings are established on the theoretical settlement curves, and subsequent measurements are plotted with reference to those initial values. The comparisons show that the calculated settlements agree reasonably well with the measured values.

2.5.4.10.4 Lateral Earth Pressures Based on anticipated groundwater elevations described in Subsection 2.5.4.6, all substructures were designed to resist full hydrostatic groundwater pressure at all levels below elevation 730 feet. All mat foundations established below elevation 730 feet were designed to resist hydrostatic uplift pressures. Subsurface walls were designed to resist lateral pressures induced by soil and groundwater under both static and dynamic loading conditions. The analysis for lateral stability of the power station, under static and dynamic loading conditions was performed for the following situation: Only Unit 1 constructed with an open excavation for Unit 2. Adequate factors of safety for lateral stability exist for the above condition (Q&R 241.20). The values of the lateral earth pressures due to the soil, water, surcharge, and dynamic loads were determined using the following equations:

SOIL Ps = Kwh, for h < h1 (2.5-5) Ps = Kwh1 + K (s - H) (h - h1), for h > h1 WATER Pw = 0, for h < h1 (2.5-6)

CPS/USAR CHAPTER 02 2.5-74 REV. 11, JANUARY 2005 Pw = H (h - h1), for h > h1 SURCHARGE Pc = KS (2.5-6) DYNAMIC INCREMENT Pe = Kh (Ho - h) w for h < h1 (2.5-8) Pe = Kh (Ho - h) w for h > h1 (2.5-9) where: H = height of wall (ft) h = distance from grade to the point where pressure is to be determined (ft) h1 = distance from final grade to the groundwater table (ft)

K = horizontal earthquake acceleration/gravity (.10 OBE, .25 SSE)

Pc = lateral pressure due to surcharge (ksf) Pe = lateral pressure due to earthquake at a distance h below grade (ksf) Ps = lateral pressure due to soil at a distance h below grade (ksf)

Pw = lateral pressure due to water at a distance h below grade (ksf)

S = surcharge loading, .5 ksf for construction, 1.0 ksf for E-70, .3 ksf for H-20 H = density of water (.0624 kcf) s = density of saturated soil (.1374 kcf) w = density of wet soil (.132 kcf) (a) The static coefficient of lateral earth pressure of 0.47 was calculated using an angle of internal friction of 32

°. This is equal to the at-rest earth pressure coefficient for cohesive (clayey silt and silty clay till) backfill used at CPS. Granular backfill has also been used for backfill around Category I structure walls. The at-rest earth pressure coefficient for granular backfill was calculated using an angle of internal friction of 38

°. The static coefficient of lateral earth pressure for the granular backfill is 0.38 which is less than the value (0.47) used and therefore is conservative. (b) The dynamic water pressure is considered in combination with dynamic soil pressure and is accounted for by using density of saturated soil in Equation 2.5-9, Subsection

2.5.4.10.4.

CPS/USAR CHAPTER 02 2.5-75 REV. 11, JANUARY 2005 (c) The lateral earth pressures due to soil, water, and dynamic loads are considered in the design by using equations given in Subsection 2.5.4.10.4. The plot showing these lateral earth pressures is shown in Figure 2.5-492. Soil-Structure Interaction Analysis as described in Subsection 3.7.2 does not give lateral pressures. If the model synthesis approach used for the analysis, only the modal properties (i.e., frequencies, mode shapes, modal damping values, and participation factors) are extracted from the soil model. Any other responses since they are not of interest for the interaction analysis (Q&R 241.7). 2.5.4.11 Design Criteria The design criteria used in the design of Seismic Category I structures are discussed in the following subsections: a. liquefaction potential, Subsection 2.5.4.8; b. bearing capacity, Subsection 2.5.4.10.2;

c. settlement, Subsection 2.5.4.10.3;

d. static slope stability, Subsection 2.5.5.2.3; and

e. dynamic slope stability, Subsection 2.5.5.2.4. 2.5.4.12 Techniques to Improve Subsurface Conditions Localized areas and pockets of loose granular materials were encountered in the base of the excavations for some of the Category I structures. These materials were either compacted and tested or removed. Depressions created by these minor excavations were filled with a flyash backfill mixture. These operations are discussed in Subsections 2.5.4.5.1.4, 2.5.4.5.2.4, and

2.5.4.5.3.4. 2.5.4.13 Subsurface Instrumentation An instrumentation system of settlement points was established at various locations throughout the buildings at the main plant site. Figure 2.5-382 shows the location of these devices. Each settlement point consists of a brass monument imbedded in the rough concrete. Reference points were located around the structures from which easy accessibility could be

made. Most settlement points were measured at a minimum frequency of once every four calendar months until the plant settlement was considered stabilized. The graphical plots of measurement readings recorded versus the time for all settlement points are shown in Figure

2.5-438. 2.5.4.14 Construction Notes Instead of using either cohesive or granular backfill in small, confined areas, a flyash mixture was used. The flyash backfill was a mixture of cement, flyash, sand, and water mixed to the following approximately proportions determined by weight per cubic yard: cement - 50 to 100 lb; CPS/USAR CHAPTER 02 2.5-76 REV. 11, JANUARY 2005 flyash - 450 lb; sand - 2800 lb; and, water - 425 lb. The material was mixed in a central concrete batching plant, transported by mixer trucks, and placed in the designated areas so as not to permit sedimentation. The material was tested in place following a curing time long enough to give the material a strength which yields a maximum deflection of 0.25 inches under an applied pressure of 50 psi. A total of 237 in-place strength tests were performed for the acceptance of the flyash mixture backfill placed beneath and around the structures including the main power block, screenhouse, outlet structure, and the pipelines. The minimum load used for the test was approximately 55 psi. The maximum deflection recorded was 0.194 inches.

Therefore, the tests performed on the bash were acceptable. The use of the flyash mixture backfill was restricted to areas beneath and around piping and for dental work on the subgrade and in confined areas. The degree of compaction of either granular or cohesive materials in these areas by conventional means could not be guaranteed due to the size awkwardness of the compaction equipment. Also, the ease of placement as well as the lack of any compaction requirements for the flyash backfill made its use more practical in these small and confined areas. 2.5.4.14.1 Main Plant The base of the excavation for the main plant was overexcavated in several areas because the inplace density of the subgrade did not meet the requirements as stated in Subsection 2.5.4.5.1.4. Depressions created in these areas were backfilled with a flyash mixture as described in Subsection 2.5.4.5.1.4. Structural backfill for the main plant structures consisted of Type B and B1 granular material. This material was utilized along with the Type A and C cohesive materials. Also, the flyash backfill was used around some piping in the main plant. These backfill operations are discussed in Subsections 2.5.4.5.1.5 and 2.5.4.5.1.6. The proposed borrow area for the Type B structural fill for the main plant was not used as the source of the granular fill. A new borrow area was developed as discussed in Subsection

2.5.4.5.1.5. 2.5.4.14.2 Circulating Water Screen House A flyash backfill mixture was used as backfill material around piping and for dental work in filling in sumps. This is discussed in Subsection 2.5.4.5.2.5. 2.5.4.14.3 SSWS Outlet Structure and Pipelines The subgrade beneath the SSWS outlet structures was overexcavated due to the presence of soft material at the proposed foundation elevation. Type B granular material was used as structural fill beneath the structure to attain the design foundation elevation. This is discussed in Subsections 2.5.4.5.3.2 and 2.5.4.5.3.5. A flyash mixture was used as fill material around the SSWS outlet structure as well as around the SSWS pipelines. This is discussed in Subsection 2.5.4.5.3.5.

CPS/USAR CHAPTER 02 2.5-77 REV. 13, JANUARY 2009 2.5.4.14.4 Main Plant Excavation The excavation for Unit 2, adjacent to the plant east side of Unit 1, will remain open as shown on Figure 2.5-484A. An administration building has been constructed in the Unit 2 excavation, as shown on Figure 2.5-484B. The slopes of the backfill adjacent to the Unit 1 structures will be graded as illustrated by Sections G-G and H-H. After grading, a revetment composed of a grout intrusion blanket will be placed on these slopes to protect them against erosion due to runoff. A berm, consisting of compacted Type A cohesive material or un-reinforced concrete will be placed around the perimeter of the excavation to divert flood water runoff from entering the excavation. No berm will be placed across the construction ramp. Figure 2.5-484A also illustrates the drainage system that is utilized to drain any precipitation that enters the excavation. As shown on Section A, a flap gate will be used to prevent a backflow from the lake into the excavation if the lake level rises above the invert elevation of the drainage system. In addition to the drainage system, all openings in the Unit 1 building below grade level that are exposed in the Unit 2 excavation will be closed and waterproofed. 2.5.5 Stability of Slopes 2.5.5.1 Slope Characteristics The alignment of the cooling lake main dam is shown in Figure 2.4-13 and a typical cross section showing the as-built details is presented in Figure 2.4-14. The layout of the ultimate heat sink and the as-built features of the submerged earthfill are shown in Figures 2.5-384, 2.5-385, 2.5-386, and 2.4-24. There are no natural or manmade slopes in the immediate vicinity of the plant whose failure would adversely affect the safety of the power plant. 2.5.5.1.1 Main Dam The main dam was constructed with side slopes of 3:1 (horizontal to vertical) for both the upstream and downstream sides. The dam section is a homogeneous embankment constructed using the silty clay materials of the Wisconsinan glacial till of the Wedron Formation (Type A) obtained from the borrow areas. A downstream drainage blanket and a ditch were provided to drain away the small seepage that will occur through the dam. The drainage blanket will also help lower the phreatic line and prevent seepage from emerging on the downstream slope of the dam, thus preventing the softening and erosion of the downstream slopes. A typical cross section is illustrated in Figure 2.4-14. The upstream slope of the dam is protected against destructive wave action by means of rock riprap and bedding material. The riprap design and other slope protection for the main dam is discussed in Subsections 2.4.8 and 2.5.6.4.2.2. The average soil properties, based on laboratory tests performed on the borrow materials, are summarized in Table 2.5-52. The subsurface exploration and local geologic features at the main dam site are discussed in Subsections 2.5.4.3, 2.5.4.4, and 2.5.6.2.2. 2.5.5.1.2 Ultimate Heat Sink The subsurface soil, rock, and groundwater conditions at the ultimate heat sink were explored by drilling 60 test borings (H-borings) at the locations indicated in Figure 2.5-16. Twenty-seven borings were drilled for the submerged dam, its abutments, and vicinity; twenty-three, for the CPS/USAR CHAPTER 02 2.5-78 REV. 11, JANUARY 2005 excavated and natural slopes; six, for the baffle dike and its abutment; and four, for the screen house and outlet structures. The static and dynamic properties of the foundation soils and the materials used in the construction of the submerged dam were evaluated through a comprehensive drilling and testing program described in Subsections 2.5.4.2 and 2.5.5.3.2. This program also included evaluation of the properties of the soils that form the slopes of the cut areas of the heat sink. The as-built features of the ultimate heat sink submerged dam and baffle dike are shown in cross sections depicted in Figures 2.5-384 and 2.4-24.

The submerged dam and the baffle dike were constructed using the Wisconsinan Glacial Till of the Wedron Formation (Type A material) removed from borrow areas located south of the heat sink and shown in Figure 2.5-384. Side slopes of 5:1 (horizontal to vertical) for both the upstream and downstream sides of the submerged dam and the baffle dike were constructed and were determined to be stable under static and seismic loading conditions. Slope protection for the baffle dike, the s ubmerged dam, and its abutments was provided by a soil-cement mixture placed as shown in Figures 2.5-386 and 2.4-24. The mixing and placement procedures for the soil-cement are described in Subsection 2.5.6.4.1.2. The stability of the slopes bordering the ultimate heat sink was examined by preparing ground profiles based on the original topography and modified as necessary by the aerial survey taken prior to the closing of the submerged dam and lake filling (Figure 2.5-385). Soil profiles based on the boring logs and field conditions used in the analysis are provided in Figures 2.5-387 through 2.5-392. The maximum final slope of the natural ground on the south and east sides of the heat sink below elevation 690 is 5:1 (horizontal to vertical). Final slopes for the compacted fill material for the baffle dike abutment on the east side above elevation 675 is 4:1 (horizontal to vertical). The final excavated slopes in the vicinity of the screen house and outlet structure on the east side are 5:1 (horizontal to vertical) below elevation 690 and 3.5:1 (horizontal to vertical) above elevation 690. The remaining final slopes on the east side are flatter than 3.5:1 (horizontal to vertical) above elevation 690. The natural slopes above elevation 690 on the south side are flatter than 5:1 (horizontal to vertical), except at the locations of Section Y'-Y' and G-G, as shown in Figure 2.5-385. The maximum slope of the northern natural ground above elevation 668.5 is 7:1 (horizontal to vertical). The submerged dam provides the boundary along the western section of the heat sink. The subsurface conditions in the ultimate heat sink area and adjacent to the valley walls are described in Subsections 2.5.5.2.3.2 and 2.5.6.4.1. The soil properties and other conditions considered in the stability analyses are presented in Subsection 2.5.5.2.3. The dynamic stability analysis is described in Subsection 2.5.5.2.4. Cyclic triaxial tests have been performed for determining the dynamic properties and the results are discussed in Subsection 2.5.5.2.4. At the west side of the ultimate heat sink, where the submerged dam was constructed, the ground surface elevation ranges from 670 to 673 feet in the floodplain, as revealed by the elevations of Borings H-36, H-35, H-7, H-5, H-4, and H-34. Boring H-37 (693.9 feet elevation) is located at the south abutment on the south valley wall; Boring H-51 (690.3 feet elevation) approximately at the location of the north abutment on the north side. The locations of these borings can be found in Figure 2.5-16.

CPS/USAR CHAPTER 02 2.5-79 REV. 11, JANUARY 2005 The subgrade preparation for the submerged dam removed all unsuitable soils revealed by the borings. The embankment was constructed on the Illinoian glacial till of the unaltered Glasford

Formation. The construction of the embankment required the excavation of the soils at the abutments of the proposed embankment as shown in Figure 2.5-386. The soils encountered in the abutments are discussed in Subsection 2.5.6.3.1. These excavated areas were backfilled with the Type A cohesive fill material obtained from the designated borrow areas.

A slurry trench was constructed between the diversion channel for the North Fork and the areas that were excavated for the baffle dike and the submerged dam to minimize seepage into the excavations. A series of sumps were dug along the outer perimeters of these excavations to collect the minor seepage that occurred. Pumps were used as required to drain the collected water in the sumps. The groundwater and seepage conditions used in the stability analyses are discussed in Subsection 2.5.5.2.3.3. 2.5.5.2 Design Criteria and Analyses 2.5.5.2.1 Main Dam The main dam is designed so as to be stable under all conditions of reservoir operation. The stability of the slopes was investigated under the following loading conditions with the factors of safety so stated: a. end of construction, 1.49 b. steady seepage with normal storage pool (elevation 690.0 feet), 2.06

c. steady seepage with normal storage pool (elevation 690.0 feet) plus 0.1g horizontal earthquake force, 1.44 d. steady seepage with maximum storage pool (elevation 708.0 feet), tailwater at elevation 675.0 feet, 2.02 e. and, sudden drawdown from normal storage pool (elevation 690.0 feet), 1.50. Figures 2.5-393 through 2.5-397 illustrate the failure surfaces with the minimum factors of safety as determined by the computer program, BISHOP, utilizing the simplified Bishop method. The BISHOP computer program is described in Appendix C. In the simplified Bishop method, the failure surface is assumed to be an arc of a circle. The factor of safety is defined as the ratio of the moment about the center of the available resisting forces along the failure arc to the moment tending to cause sliding. A minimum factor of safety of 1.3, 1.5, 1.5, and 1.0, respectively, is required for the loading conditions listed previously in Items a, b, d, and e. Under seismic loading conditions, a minimum factor of safety of 1.0 is required for the loading condition listed previously in Item c.

CPS/USAR CHAPTER 02 2.5-80 REV. 11, JANUARY 2005 The analyses indicated that a side slope of 3:1 (horizontal to vertical) is adequate for the upstream and downstream slopes of the main dam to ensure its stability under both static and seismic loading conditions, with conservative factors of safety. In the downstream slope of the main dam, the piezometric levels in the embankment and foundation materials will increase with time as steady state seepage is established. As this occurs, the pore pressures increase and the effective stresses decrease.

In the analysis, the effective strength parameters for the soils are used; however, the pore pressures are also considered in determining the effective stresses in the dam when computing the soil strength available to resist failure. The filling of the reservoir causes no change in consolidation stress in the downstream slope of the dam. Therefore, the critical condition will occur when the seepage line is established and the greatest increase in pore pressure and corresponding reduction in effective strength has occurred. The soil properties used in the analysis for the steady state condition were conservatively chosen from the consolidated, undrained triaxial tests. If the embankment has not consolidated under its own weight such that consolidated, undrained triaxial tests are applicable, the unconsolidated, undrained test results should be used. These properties were used for the end of construction analysis. This condition is the same as considered for the end of construction analysis where the unconsolidated, undrained soil properties were used. The Salt Creek alluvium in the foundation and the sand drainage blanket in the downstream slope being sand materials should provide drainage in the embankment (Q&R 241.11). 2.5.5.2.2 Submerged Dam and Excavated Slopes of Ultimate Heat Sink The stability of the slopes of the submerged dam and the excavated slopes of the cut areas of the ultimate heat sink were investigated by applying the Seismic Category I design criteria. These analyses are based on the static and dynamic properties of the fill material, subsoil, and the bedrock developed through exploratory borings and laboratory tests.

In this analysis, the following loading conditions were considered: a. submerged dam - 1. end of construction,

2. submerged condition under normal cooling lake elevation (690.0 feet),

3. sudden drawdown condition as a result of breach of the main dam, maximum storage pool elevation (675.0 feet) in the heat sink, and complete and rapid drawdown on the downstream side, and 4. same loading conditions as in Item 2 plus safe shutdown earthquake force; and b. excavated slopes -

CPS/USAR CHAPTER 02 2.5-81 REV. 11, JANUARY 2005 1. end of construction, 2. submerged condition under normal cooling lake elevation (690.0 feet) in the heat sink plus safe shutdown earthquake force, and 3. sudden drawdown condition from normal cooling lake elevation (690.0 feet) to maximum storage pool elevation (675.0 feet) in the heat sink. The static stability of the slopes was investigated using the simplified Bishop method, implemented in the SLOPE and BISHOP computer programs, and based on the criteria presented in Subsection 2.5.5.2.3. Dynamic analyses were performed to evaluate the stability of the slopes under seismic loading conditions. These analyses were accomplished with finite-element models which consider nonlinear stress-strain and damping relations and cyclic strength characteristics of the soil. The details of these analyses are given in Subsection 2.5.5.2.4. The upstream and downstream slopes of the earthfill are protected against erosion by soil cement. The design of the soil cement slope protection is discussed in Subsection 2.5.6.4.1.2. The natural slopes bordering the ultimate heat sink on the south side generally have a maximum grade of 5:1 (horizontal to vertical), as shown in Figure 2.5-385, Sections E-E, F-F, G-G, and Y-Y. However, in the vicinity of Section G-G, the distance between the bottom of the heat sink and the steep valley wall is approximately 160 feet. At this location, the natural slope from elevation 675 to 690 is approximately 1.5:1 (horizontal to vertical). If a postulated slope failure did occur in this area and the final slope was 30:1 (horizontal to vertical), the material would only fill at the most the inlet as shown in Figure 2.5-384. The sloughed material would extend approximately 130 feet from the toe of the steep slope and would not travel into the heat sink bottom. Therefore, this material would still not reduce the capacity of the heat sink or obstruct the flow path of the circulating water. Also, in the vicinity of Section Y-Y the south side, the slopes are 3:1 (horizontal to vertical) between elevations 690 and 736. This slope is considered in the stability analysis in Subsection 2.5.5.2.3. The normal pool of the cooling lake will be at elevation 690.0 feet. Under this condition, the slopes of the uplands below elevation 690.0 feet will be completely submerged. The seepage and the zone of saturation will be dependent on the surrounding groundwater conditions. The stability of the natural slope in the immediate vicinity of the outlet structure was investigated based on the strength characteristics of the soils under saturated conditions. The static and dynamic stability of the natural slopes bordering the ultimate heat sink were investigated based on the results of the exploratory borings and laboratory tests that were performed. Possible loss of shear strength due to softening of the materials under submerged conditions was also considered in the analysis. The natural steep slopes bordering the ultimate heat sink were graded to ensure slope stability under all loading conditions. This scheme also provides protection against localized sloughing of the slopes. Figure 2.5-385 shows the final graded slopes in the heat sink area. The intent of Figure 2.5-384 was not to show the limits of the sloughed material but to indicate how the limits would relate to the heat sink configuration.

CPS/USAR CHAPTER 02 2.5-82 REV. 11, JANUARY 2005 At section G-G the maximum slope of 15:1 exists; however there is a cove or inlet along the heat sink perimeter at this location, and the toe of this slope is 160 feet from the bottom of the heat sink. If this slope is postulated to fail and slough to a 30:1 slope, the material would move 130 feet and still be 30 feet from the bottom of the heat sink. This would at most only fill the inlet or cove and not encroach on the heat sink perimeter shown in Figure 2.5-384 (Q&R 241.12). 2.5.5.2.3 Static Slope Stability Analyses Stability analyses were performed using the simplified Bishop method implemented in the SLOPE and BISHOP computer programs. The descriptions of these computer programs are given in Appendix C. Three critical slopes in the ultimate heat sink were analyzed as follows: a. The as-built, excavated slope with the excavated toe, Section X-X, at the east side of the ultimate heat sink (UHS) and in the vicinity of the outlet structure. b. The as-built, excavated slope with the excavated toe, Section Y-Y, at the south side of the ultimate heat sink. Profile used for the analysis was obtained from Borings H-23 and H-40 as shown in Figure 2.5-390. c. The as-built, excavated slope in the southeast corner, Section H-H, includes fill material placed in an existing valley. The profile used in the analysis was obtained from field data and Borings H-30, P-61, and P-61A. These three critical sections were selected based on an evaluation of the variables that were relevant to the stability of natural slopes. The variables and loading conditions considered in the analyses are presented in Subsections 2.5.5.2.3.1 through 2.5.5.2.3.8. 2.5.5.2.3.1 Slope Configuration Several combinations of slope height and slope inclinations are shown in typical cross sections through the natural slopes bordering the ultimate heat sink, Figure 2.5-385. The location of these cross sections are identified in Figure 2.5-384. Sections J-J, K-K, H-H, and X-X represent the ground profiles on the east side of the ultimate heat sink, in the vicinity of Seismic Category I structures. Sections E-E, F-F, G-G, and Y-Y represent the ground profiles on the south side of the ultimate heat sink. The variations of slope height above elevation 690.0 feet on the east and south sides, the maximum slope inclination, and the percentage of height where the maximum inclination occurs, are as follows: SECTION SLOPE HEIGHT (feet) MAXIMUM SLOPE PERCENTAGE OF HEIGHT East Side J-J 45 3.5:1 100 K-K 45 4.0:1 100 CPS/USAR CHAPTER 02 2.5-83 REV. 11, JANUARY 2005 X-X 45 3.5:1 100 H-H 45 3.5:1 100 South Side E-E 45 5.0:1 89 F-F 32 6.0:1 31 G-G 22 4.0:1 73 Y-Y 36 3.0:1 55 The height of the excavated slopes (7:1, horizontal to vertical) at the north side ranges from 5 to 20 feet. 2.5.5.2.3.2 Subsurface Conditions An extensive program of soil exploration described in Subsections 2.5.4.3 and 2.5.5.3.2 was undertaken to determine the subsurface conditions in the ultimate heat sink. The soil profiles determined by these investigations are shown in Figures 2.5-387 through 2.5-392. The locations of these profiles are shown in Figure 2.5-384. Sections P-P and Q-Q show soil profiles through the ultimate heat sink including the baffle dike. Sections X-X, Y-Y, R-R, and S-S show soil profiles along sections through the east, south, and north sides of the ultimate heat sink, respectively. A representative soil profile for the east and south sides of the ultimate heat sink was developed for use in the stability analysis. The stratigraphic units and elevations shown in the following table were defined based on the soil profiles of Sections P-P, Q-Q, X-X, and Y-Y (Figures 2.5-387 through 2.5-390) and site geologic profiles (Figures 2.5-275, 2.5-279, and 2.5-284), as well as actual field conditions encountered. Elevation (feet)

Stratigraphic Unit From To Richland Loess 740 730 Wisconsinan Glacial Till of the Wedron Formation 730 695 Interglacial Zone of the weathered Glasford Formation and Robein Silt where present 695 680 Sand Layer of the upper Glasford

Formation 680 675 Illinoian Glacial Till of the unaltered

Glasford Formation 675 570 Lacustrine Deposits of the Banner

Formation 570 560 Pre-Illinoian Glacial Till of the

Banner Formation 560 480 CPS/USAR CHAPTER 02 2.5-84 REV. 11, JANUARY 2005 Bedrock 480 The Richland loess consists of medium to stiff clayey silt or silty clay with trace of fine sand, weathered. The Wisconsinan glacial till of the Wedron Formation consists of stiff to hard clayey silt and silty clay with sand and gravel-sized particles randomly interspersed throughout the matrix or in well sorted lenses of varying thickness. The interglacial zone, consisting of the Robein Silt and the weathered Glasford Formation, is composed of stiff to very stiff silt or silty clay, clayey silt, sand, and gravel. The Illinoian glacial till of the unaltered Glasford Formation is a very hard and dense material and consists of clayey silt with occasional interspersed sand and gravel-sized particles and pockets of fine to coarse sand. The lacustrine deposits of the Banner Formation contain stiff to hard clayey silt or silt and clay with sand and gravel. The pre-Illinoian glacial till of the Banner Formation is composed of silty clay and clayey silt with some interspersed sand and gravel. Bedrock consists of interbedded layers of limestone, shale, and siltstone. The conservative design effective soil parameters assigned to each stratigraphic unit are described in Subsection 2.5.5.2.3.7. The stratigraphic units down to and including the Illinoian glacial till of the unaltered Glasford Formation form the natural slope. The remaining stratigraphic units are below the toe elevation of the natural slopes. The slope configuration above toe elevation (height and inclination) is the governing factor for the stability analysis. Since Section X-X (east side) and Y-Y (south side) had the maximum height and the steepest inclination, those sections were considered the most critical. Section H-H (east side) includes fill material placed above elevation 690 and was considered a separate case to be analyzed. The soil profiles (Figures 2.5-391 and 2.5-392) at the north side of the UHS consist of alluvial materials of the Henry Formation overlying the Illinoian glacial till. The total alluvial fill, including the topsoil, ranges in thickness from 12 feet to 23 feet. These alluvial deposits include an upper zone, 5 to 10 feet in thickness composed of soft to medium silty clay and/or clayey silt. 2.5.5.2.3.3 Groundwater Table The groundwater table was established from the elevation shown in geologic Section B-B' (Figure 2.5-275).

CPS/USAR CHAPTER 02 2.5-85 REV. 11, JANUARY 2005 2.5.5.2.3.4 Water Elevation in the Ultimate Heat Sink The elevation of the water surface in the ultimate heat sink and/or cooling lake was established at either elevation 675 or 690, depending on the condition being considered in the stability

analyses. 2.5.5.2.3.5 Earthquake Forces Analyses were performed with and without earthquake forces. When earthquake forces were used for pseudo-static cases, a horizontal ground acceleration of 0.25g was used. 2.5.5.2.3.6 Stability Conditions Three different conditions were considered in the analyses of the stability of the slopes: a. end of construction, b. full cooling lake, and c. empty cooling lake (rapid drawdown). End of construction simulates the conditions at the time when the ultimate heat sink (UHS) is completed and the slopes are finished with the excavated toe. There will be no water contained

in the UHS. Full cooling lake simulates the conditions existing during normal operation of the station with water surface established at elevation 690 feet. Empty cooling lake simulates the conditions existing in the event that the main dam fails and the water contained in the cooling lake is lost, rapid drawdown.

For end of construction and full cooling lake conditions, a steady seepage of groundwater toward the face of the slope was established with a piezometric surface of top flow line. This line defines submergence of the slope where a body of water exists. The groundwater surface is represented by a piezometric surface which calculates the buoyant force on the overburden column of soil. This is referenced to the soil or soils which are affected by this pore pressure. For the end of construction condition, the piezometric surface was established from the groundwater table shown in geologic Section B-B' (Figure 2.5-275). For full cooling lake condition, the water surface was established at elevation 690 in the ultimate heat sink and the groundwater surface in the slope (for these two conditions, see Figures 2.5-398 and 2.5-399 for Section X-X, Figures 2.5-400 and 2.5-401 for Section H-H, and Figures 2.5-489 and 2.5-490 for Section Y-Y). The empty cooling lake condition is represented by a drawdown case which is the lowering of the water level against the slope. The initial water level and steady state after drawdown level are defined by two piezometric surfaces, 2 and 1, respectively. The pore pressure calculations are a function of piezometric surfaces 1 and 2 (see Figure 2.5-402 for Section X-X and Figure 2.5-491 for Section Y-Y).

CPS/USAR CHAPTER 02 2.5-86 REV. 11, JANUARY 2005 For this case, the initial water level is the lowering of the piezometric surface established for full cooling lake condition against the slope and the drawdown level is 675 feet, the maximum storage pool elevation in the ultimate heat sink. For Section H-H, the piezometric surface was defined by the anticipated water level immediately following the loss of the cooling lake down to elevation 675 feet as shown in Figure 2.5-403.

Static and pseudo-static analyses were performed for end of construction and full cooling lake cases; only static analyses for the empty cooling lake case. Effective stress parameters were utilized in all cases. The end of construction condition was not considered for Section H-H because the fill material had been in place for some time prior to the stability analysis of the section and no sloughing of material was noticed. 2.5.5.2.3.7 Selection of Soil Parameters The effective strength parameters (cohesion and angle of internal friction) for the soil deposits comprising the slopes were obtained from the Mohr's circles plots using the results of consolidated undrained triaxial tests with measurements of pore pressure. These consolidated undrained triaxial test data are shown in Tables 2.5-10, 2.5-11, and 2.5-13, and the Mohr's circles are presented in Figures 2.5-404 through 2.5-408. The strength parameters for the loess and the sand layer in the upper Glasford Formation were selected from published data showing correlations between number of blow counts and angle of internal friction. The following are the correlations used to obtain the strength parameters.

Soil Type Average N-Value (blows/foot) Relative Density Angle of Friction Loess 28 NA 20° Sand 44 75%

38° The static moduli of elasticity were calculated using the stress-strain curves of the triaxial tests, at the strain corresponding to one-half of the peak deviator stress. The densities shown in Table 2.5-53 are averages of natural densities obtained from numerous samples taken from the borings. The Poisson's ratios are assumed values taken from published data. The soil properties used in the analysis for Section H-H are based on the borings in that area and are provided in Figures 2.5-400, 2.5-401 and 2.5-403. The design effective strength parameters are shown in Table 2.5-53. Cohesion and friction values correspond approximately to the minimum of the strength parameters of all soils within each stratigraphic unit, based on the data presented in Figures 2.5-404 through 2.5-408. The blow counts corresponding to the samples used in triaxial compression tests fall within the range of the minimum blow counts for the soil layers within each stratigraphic unit.

Consequently, the strength parameters are representative of the weak soils with low blow counts and low density. Blow counts less than those shown in the data of Figures 2.5-404 through 2.5-408 are found in the Wisconsinan till at Boring P-38, (elevation 730), and in the Illinoian till at Boring P-44, (elevation 680 feet). Since the excavation for the construction of the CPS/USAR CHAPTER 02 2.5-87 REV. 11, JANUARY 2005 station, screenhouse, and the vicinity of the outlet structure is removed from these soft soils, strength parameters for these soils were not determined. Based on the conservatism used in the selection of strength parameters, the same cohesion and friction values were assigned to the Wisconsinan till of the Wedron Formation and interglacial cohesive coils of the weathered Glasford Formation. Conservative parameters for the Illinoian till of the unaltered Glasford Formation were chosen for stability analysis (42

° angle of internal friction and 1400 lb/ft 2 cohesion). These parameters were selected based on the Mohr's circles plot shown in Figure 2.5-406. The same properties were assigned to the pre-Illinoian lacustrine or glacial till due to similarity in soil properties. The soil profile for Section X-X of the Ultimate Heat Sink is shown in Figure 2.5-389. This figure shows the sand layer underlying the Interglacial Zone which is designated as Soil No. 4 in the model used for the stability analysis (Figure 2.5-398). As discussed in Subsection 2.5.5.2.3.9, this sand layer belongs to the upper Glasford Formation and has an average N-value of 44 blows per foot which corresponds to 75% relative density. The strength parameters for this material were selected as ø equal to 38

°, c equal to 0 from correlations between number of blow counts and angle of internal friction (Subsection 2.5.5.2.3.7). The blow counts in the range 2, 4, 5 in borings P-12 and P-8 are for clayey silt (ML) and silty clay (CL) which overlay the sand (SP) soils. The sand materials in these two borings had N-values of 23 and 25 relative densities of 91% and 95% (Table 2.5-54). The clays and silts in borings P-8 and P-12 lie above the elevation 668.5 and were removed during excavation of the heat sink and toe of slope. The material below the excavation is sand with relative densities of 75% or greater for which the equal to 38

°, C equal to 0 strength parameter are applicable. Published data (H. F. Winterkorn and H. Y. Fang, Foundation Engineering Handbook, Van Nostrand Reinhold Co., 1975) show that for dense sand (N = 30 to 50, relative density = 65% to 85%). The range of values is 36

° to 41° according to Peck and 40

° to 45° according to Mayerhoff (Table 2.43 p. 117). A compilation of many relationships between relative density vs. friction angle for cohesionless soils (Figure 7.26, Winterkorn and Fang, p. 263) also shows that for relative densities of 75% and greater, the angle of internal friction is = 38° or greater (Q&R 241.13). The slope stability analysis at Section H-H of the Ultimate Heat Sink as shown in Figure 2.5-400 used shear strength parameters derived from samples of the Wisconsinan Till compacted to 90% modified optimum density as given in Table 2.5-13. During construction of the Main Dam, the embankment was partially completed and left to stand during the winter of 1976-77. At the resumption of work in the spring of 1977, samples were taken from the zone of frost penetration to determine what effects the frost had on the soil properties. These soil properties were used in the Main Dam analysis and are shown in Table 2.5-52. Subsection 2.5.6.9 describes this testing program. The use of the soil properties for frost affected soil was to show that an adequate factor of safety is available even if the reduced soil parameters were used. The fill in the Ultimate Heat Sink slope was constructed during a single construction season and material interior to the fill

was not subject to frost.

CPS/USAR CHAPTER 02 2.5-88 REV. 11, JANUARY 2005 The normal pool elevation shown in Figure 2.5-403 is intended only as a reference level from which the drawdown occurred. The water level at elevation 675 feet was used in the analysis (Q&R 241.14a). 2.5.5.2.3.8 Results of Stability Analyses The results of the stability analyses for as-built excavated slopes are presented in the following Sections (X-X, Y-Y, and H-H) and graphically in Figures 2.5-398, 2.5-399, and 2.5-402 for Section X-X, Figures 2.5-400, 2.5-401, and 2.5-403 for Section H-H, and Figures 2.5-489, 2.5-490, and 2.5-491 for Section Y-Y. The critical circle is drawn on the section the referenced to the center by the radius lines. FACTORS OF SAFETY CONDITIONS SECTION X-X SECTION Y-Y SECTION H-H End of Construction Static 2.60 2.42 NOT Pseudo-Static 1.24 1.21 ANALYZED Full Cooling Lake Static 2.33 2.15 2.32 Pseudo-Static 1.07 1.03 1.02 Empty Cooling Lake Static 1.96 2.09 2.16 The stability analysis for Section X-X considered an excavated slope configuration of a 5:1 (horizontal to vertical) slope from the heat sink bottom to elevation 690, and then a 3.5:1 (horizontal to vertical) slope up to natural ground at approximately elevation 736. The stability analysis for Section Y-Y considered an excavated slope configuration of 5:1 (horizontal to vertical) from the bottom of the heat sink to elevation 690, and then a 3:1 (horizontal to vertical) slope from elevation 690 to elevation 736. Based on the results of the static and pseudo-static stability analyses, the slopes of the ultimate heat sink are stable with an adequate factor of safety. At the north side of the ultimate heat sink, the excavated slopes are very flat (7:1, horizontal to vertical). Attachment A2.5 discusses this slope in greater detail with regard to slope stability. The submerged dam at the west side of the ultimate heat sink, with a maximum height of 21 feet and side slopes of 5:1 (horizontal to vertical), was constructed using Wisconsinan glacial till of the Wedron Formation (Type A material). Static analyses performed for the main dam having a maximum height of 57 feet and side slopes of 3:1 (horizontal to vertical) constructed with Wisconsinan till of the Wedron Formation (Type A material), show adequate factors of safety.

Based on these analyses, it is concluded that the slopes of the submerged dike are also stable under an adequate factor of safety. Finite-element analyses are presented in Subsection 2.5.5.2.4.

CPS/USAR CHAPTER 02 2.5-89 REV. 11, JANUARY 2005 2.5.5.2.3.9 Liquefaction Potential From the point of view of liquefaction, the alluvial soils of the Henry Formation on the floodplain and the sand layer in the upper Glasford Formation on the valley walls are of interest. The presence of these soils was determined by the borings located in Figure 2.5-16. The simplified procedure for evaluating soil liquefaction described by Seed and Idriss (Reference 85) has been applied to the alluvial materials which have a relatively level surface. The finite-element dynamic analysis described in Subsection 2.5.5.2.4.4 incorporates the sand layer behind the natural slopes. The alluvial soils at the east side of the ultimate heat sink were evaluated from Borings P-5, H-28, H-32 (vicinity of the screen house) and Borings P-8, P-12, and H-33 (vicinity of the outlet structure). From the results shown in Table 2.5-54, it is seen that the soils having relative densities greater than 90% are safe against liquefaction. Borings H-28, H-32 and H-33 indicated the presence of some loose alluvial material above elevation 667 feet. The excavation for the screen house and outlet structure removed this material (Figures 2.5-379 and 2.5-381).

The excavated areas around the screen house and outlet structures were backfilled with either compacted Wisconsinan till of the Wedron Formation (Types A and C material) or fly ash backfill. The extent of the backfill materials is discussed in Subsection 2.5.4.5.2 for the screen house, and Subsection 2.5.4.5.3 for the outlet structure. The alluvial material in other borings in this area have relative densities in excess of 90% with factors of safety against liquefaction ranging from 1.01 to 1.19 The sand layer in the upper Glasford Formation revealed by Borings P-9, P-13, P-37, P-48, H-30, and H-31 (east side) and Borings H-23 and H-26 (south side) has an average N-value of 44 blows per foot, which corresponds to 75% relative density. This deposit consists of clean, silty sand and gravel of varying gradation. The dynamic analysis performed for Section X-X (the most critical subsoil profile discussed in Subsection 2.5.5.2.3.2) shows that the elements of the sand layer are safe against liquefaction. This is described in Subsection 2.5.5.2.4. It is concluded, therefore, that the excavated slopes, the backfill, and the alluvial materials in the vicinity of the screen house, outlet structure, and the east side of the heat sink will maintain their integrity and not liquefy during the safe shutdown earthquake. Along the north and south sides of the periphery of the ultimate heat sink, a flow-slide method of analysis was used to determine the final slope conditions after the safe shutdown earthquake is postulated to have caused flow-type slides of the side slopes.

Attachment A2.5 discusses four postulated failure modes that could possibly reduce the capability of the UHS to function normally. To verify that adequate capacity remains in the heat sink following any local seismic activity with acceleration greater than the OBE, the bottom of the heat sink will be monitored as described in

Subsection 2.4.11.6.

CPS/USAR CHAPTER 02 2.5-90 REV. 11, JANUARY 2005 2.5.5.2.4 Dynamic Slope Stability Analysis 2.5.5.2.4.1 Method of Analysis The stability of the submerged dam and the natural slopes under seismic loading conditions was investigated using the dynamic finite element analyses which consider the strain-dependent material properties. For the dynamic analyses, the postulated safe shutdown earthquake presented in Subsection 2.5.2.6 was used. The loading conditions for which the dike and natural slope were analyzed are discussed in Subsection 2.5.5.2.2. The static and dynamic properties of the fill material and foundation soils were based on a comprehensive soil exploration and laboratory testing program as discussed in Subsection 2.5.4.2. The procedure used in evaluating the seismic stability of the slopes consists of the following steps: a. determination of the response of the dam foundation system to the compatible rock accelerations, including the evaluation of the induced shear stresses at various locations throughout the dam and the foundation material; b. representation of the irregular cycles of shear stresses induced in the dam foundation system by an equivalent number of cycles of equivalent uniform shear stresses; c. determination of the static stresses existing in the dam foundation system (prior to the rock acceleration); d. determination of the cyclic shear stresses required to cause single amplitude shear strains greater than 5 x 10

-2 in the material for conditions representative of those existing in the dam foundation system by means of appropriate cyclic load tests on representative specimens of the materials; and e. evaluation of the seismic stability of the dams by comparing the shear stresses required to cause single amplitude shear strain greater than 5 x 10

-2 with the equivalent shear stresses induced by rock accelerations. 2.5.5.2.4.2 Step-By-Step Procedure Used in Seismic Stability Evaluation The following steps were used in evaluating the seismic stability of the dike and the natural

slope: 1. Generation of a Compatible Time-History, G(t)

Synthetic accelerograms were generated for horizontal and vertical ground motions such that the response spectra of these accelerograms envelope the design response spectra. For this purpose, the N-S component of the 1940 El Centro earthquake record was suitably modified using the RSG program described in Appendix C. The normalized accelerograms are shown in Figures 2.5-409 and 2.5-410 for the horizontal and vertical ground accelerations, respectively.

The close matching of the response spectra obtained for the synthetic accelerogram with the design response spectra is demonstrated in Figures 3.7-1 through 3.7-10.

CPS/USAR CHAPTER 02 2.5-91 REV. 11, JANUARY 2005 2. Generation of a Compatible Horizontal Rock Motion, I H (t) The soil profile below the toe of the slopes was modeled as a continuous shear layer system. The compatible time-history for the horizontal ground motion as obtained in Step 1 was input at the top layer of this model. A maximum acceleration of 0.25g was used. The soil properties were defined in terms of the curves for strain dependent soil moduli and damping ratios. By inputting the compatible time-history for horizontal ground motion and iterating on the strain dependent soil properties, a compatible horizontal rock motion was obtained using the computer

program SHAKE, described in Appendix C. 3. Generation of a Compatible Vertical Rock Motion, I v (t) The vertical rock motion was obtained by exciting the SHAKE model used in Step 2 with the compatible time-history for the vertical ground motion. A maximum acceleration of 0.25g was used. The soil properties used for this step were corresponding to the properties obtained in Step 2. The vertical rock motion I v (t) thus obtained was used as input at the rock level in the two dimensional model used in Step 4. 4. Dynamic Response Analysis The horizontal rock motion I H (t) and the vertical rock motion I v (t) were used simultaneously, for the dynamic response analysis of the dike and the slope to evaluate the shear stress time-histories at various locations in the dike and slope. The response computation was performed using the finite element method of analysis. The computer program QUAD-4, described in Appendix C, was used to compute the response in the embankment foundation system. This program incorporates the use of strain-dependent modulus and damping ratio for each element of the model. The width of the model was kept large enough to ensure free field conditions at the remote boundary. Several iterations were made on the soil properties and finally, the shear stress time-history generated in each element as a result of the simultaneous horizontal and

vertical rock motion was obtained. 5. Representation of Irregular Shear Stress Time-History by Equivalent Uniform Shear Stress The procedure used to represent the irregular shear stress time-history at any element by equivalent uniform shear stress corresponding to any N c number of cycles, involves the following operations: a. sorting the peak stresses, saving only the largest stress value between each zero crossing (both positive and negative peak values are saved); b. rearranging the peak shear stresses in descending order so that the largest peak is first and the smallest is last; c. averaging of the absolute values of the positive and negative stress peaks for each cycle; and d. the cumulative average shear stress is computed for increasing number of cycles. The cumulative average value corresponds to the equivalent shear stress for the N c number of cycles for which the average has been obtained.

CPS/USAR CHAPTER 02 2.5-92 REV. 11, JANUARY 2005 6. Static Stress Analysis A knowledge of the initial static effective stress conditions is required for the evaluation of the cyclic strength of materials in the dike and the slope. For the dike, an incremental finite-element approach was used which simulates the construction of an embankment in a series of layers. The dike was divided into several horizontal layers each represented by quadrilateral elements.

During any increment of the layer, appropriate values of modulus E and Poisson's ratio were assigned to each element. After determining the stresses, E and were reevaluated for the average stress conditions during the new increment and compared with the assigned values. If a significant difference was obtained, the E and values were adjusted until a reasonable correspondence was established between the input and the computed values. This process was continued until the last layer was added. The effect of buoyancy on stresses was evaluated by using submerged unit weights for all materials in the dike and its foundation. The analysis was conducted using the computer program EMBANK, described in Appendix C. For determining stresses in the foundation of the dike and the entire natural slope, the weight of all the materials of the slope was turned on simultaneously, and the stresses were determined assuming the at-rest condition. 7. Dynamic Material Properties In order to conduct the analysis, the cyclic shear stresses required to cause single amplitude shear strains greater than 5% in all the materials for conditions representative of those existing in the slope, dike, and the foundations must be determined. The data presented in Figures 2.5-411 and 2.5-412 have been derived from dynamic triaxial compression test results on recompacted Wisconsinan till of the Wedron Formation, Type A material, conducted in accordance with the procedures described by Seed and Peacock (Reference 94). It has been assumed that the data are applicable to interglacial zone, Wisconsinan till of the Wedron Formation, compacted fill of Wisconsinan till of the Wedron Formation (Type A material) and Illinoian till of the unaltered Glasford Formation. Based on an evaluation of strength parameters given in Tables 2.5-55 and 2.5-53, respectively, it is anticipated that compacted Wisconsinan till of the Wedron Formation, Type A material, is li kely to exhibit minimum dynamic strength as compared with the strength values for other materials involved herein. Therefore, it appears to be conservative to use the dynamic properties of compacted Wisconsinan till of the Wedron Formation, Type A material, for representing the material properties of the four materials. The dynamic material properties given in Figure 2.5-413 refer to the in situ interglacial sand deposits overlying the Illinoian till, of the unaltered Glasford Formation, and were obtained from appropriate dynamic triaxial compression tests. 8. Evaluation of Seismic Stability The minumum factor of safety for various elements against local failure due to seismic loading was determined by comparing the shear stresses required to cause single amplitude shear 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 the procedure described in Step 5. A minimum value of this stress ratio greater than approximately 1.1 is generally considered to provide an ample margin of safety.

CPS/USAR CHAPTER 02 2.5-93 REV. 11, JANUARY 2005 2.5.5.2.4.3 Description of the Submerged Dike and Natural Slopes The geometry of the cross sections of the submerged dike and that of the natural slope analyzed for the seismic stability are shown in Figures 2.5-414 and 2.5-415, respectively.

Based on the height, the steepness of the slopes, the depth to bedrock and the soil data for the site, these two cross sections were considered to be the most critical sections for the dike and the east side slope of the ultimate heat sink (UHS) in the vicinity of Seismic Category I structures. These are referred to as maximum cross sections. Due to construction features, the excavated slopes in the vicinity of the screen house and outlet structure are now 5:1 (horizontal to vertical) for the slope below elevation 690 feet. Thus these slopes are flatter than the natural slope cross section used in the analysis. This flattening of the slopes will add safety to the slopes. The submerged dike has a maximum height of approximately 17 feet with a side slope of 5:1 (horizontal to vertical) for both the upstream and downstream slopes. It has been constructed using the Wisconsinan glacial till of the Wedron Formation, Type A material, excavated from borrow areas. The maximum cross section of the submerged dike and the soil layering under the dike down to the bedrock shown in Figure 2.5-414 is based on Boring H-6. For the soils forming the natural slope and the foundation of the submerged dike, the material properties are based on laboratory test results, field measurements, and published data. 2.5.5.2.4.4 Seismic Stability Evaluation of Submerged Dike The maximum cross section of the submerged dike is shown in Figure 2.5-414. The soil profile of the material under the dike is modeled as a continuous shear layer system as shown in Figure 2.5-416. The strain dependent shear moduli and damping ratio values for all the materials of the dike and foundation are given in Table 2.5-48. To conduct the dynamic response analysis, a finite element model of the dike, shown in Figure 2.5-417, was prepared. This model was excited at its base with the horizontal and vertical rock motions, I H (t) and I v (t). A typical stress time-history induced in an element is presented in Figure 2.5-418. The procedure used for representing the irregular shear stress time-history by equivalent uniform shear stress has been illustrated in Table 2.5-64. The soil properties used in the static stress analysis of the dike are given in Table 2.5-55. The finite element model used for this analysis is similar to that being used for the dynamic analysis; however, it was necessary to use only half the model for static analysis because of the symmetry. The boundary conditions at the line of symmetry were such that only vertical displacements were permitted along this boundary of the model. To evaluate the factors of safety for various elements in the dike, the shear stresses required to cause 5% single amplitude shear strain in the elements were compared with the shear stresses induced in the elements due to earthquake (Step 8 in Subsection 2.5.5.2.4.2). Such a comparison is presented in Table 2.5-56. It will be seen from this table that the minimum factor of safety against development of 5% single amplitude shear strain exceeds the minimum required factor of safety of 1.1 for all the elements analyzed. 2.5.5.2.4.5 Seismic Stability Evaluation of Natural Slope The maximum cross section for the natural slope, as used in the analysis, is shown in Figure 2.5-415. In order to obtain the horizontal and vertical rock motion, Steps 2 and 3 in Subsection 2.5.5.2.4.2 were used. The continuous shear layer model is shown in Figure 2.5-419. The CPS/USAR CHAPTER 02 2.5-94 REV. 11, JANUARY 2005 variation of shear moduli and damping ratio values with strain for the materials in the slope and its foundation is given in Table 2.5-48. The dynamic response analysis was conducted using the finite element model shown in Figure 2.5-420. The irregular shear stress time-histories obtained for various elements of the model were represented by equivalent and uniform shear stresses of equivalent corresponding number of cycles using the procedures described in Step 5 in Subsection 2.5.5.2.4.2. The static stresses in various elements were obtained using the procedure described in Step 6 of Subsection 2.5.5.2.4.2. The soil properties used in the analysis are given in Table 2.5-53. The evaluation of seismic stability was done by determining the local factor of safety for various elements. The results are summarized in Table 2.5-57. It can be determined from this table that the minim um factor of safety against developement of 5% single amplitude shear strain in the natural slope exceeds the minimum required value of 1.1 for all elements analyzed. 2.5.5.2.4.6 Seismic Stability Evaluation of Submerged Dike and Natural Slope for a New Madrid Type Event In addition to evaluating the effect of safe shutdown earthquake, the stability of the submerged dam and the natural slopes under seismic loading conditions was also investigated for a New Madrid type event, as specified in Reference 116. A 90 seconds-long time history was developed by modifying the April 1949 Western-Washington earthquake recorded at Olympia (Washington highway test lab, N86E component) to envelope the four-response spectra due to New Madrid type event shown in Figures 35 through 38 of Reference 116. This time history was then used for the stability analysis. The procedure used for stability analyses of earth-structures is essentially similar to that described in Subsection 2.5.5.2.4.1. However, only the horizontal component of the ground motion has been considered and the motion is applied at the ground surface. The results are summarized in Tables 2.5-58 and 2.5-59. It can be seen from these tables that the minimum factors of safety for all the analyzed elements of the dike as well as of the natural slope exceed the minimum required safety factor of 1.1. 2.5.5.2.4.7 Conclusion Based on the finite elements method of analysis and using the procedure described, it can be concluded that the minimum local factor of safety in the cross sections of the submerged dike as well as of the natural slope exceed the allowable minimum factor of safety for the embankment and the facilities surrounding the ultimate heat sink will maintain their integrity under the safe shutdown earthquake and a New Madrid type event. 2.5.5.3 Logs of Borings 2.5.5.3.1 Main Dam Borings in the borrow area for the main dam are located in Figure 2.5-272. The logs of the borings are presented in Figures 2.5-129 through 2.5-144. Laboratory testing was performed on representative bulk samples of the borrow material. Results of the tests are presented in the following tables and figures: a. optimum moisture contents: Table 2.5-52; CPS/USAR CHAPTER 02 2.5-95 REV. 11, JANUARY 2005 b. particle size analyses of placed material: Figure 2.5-421; c. field compaction tests: Figure 2.5-422;

d. consolidation tests: Figures 2.5-332, 2.5-336, and 2.5-337;

e. triaxial compression tests: Tables 2.5-12 and 2.5-13; f. unconfined compression tests: Tables 2.5-8 and 2.5-9; g. Atterberg limit of placed material: Table 2.5-52; and

h. permeability tests: Table 2.5-33. A summary of the properties of the materials used to construct the main dam is presented in Table 2.5-52. 2.5.5.3.2 Ultimate Heat Sink Borings in the borrow areas and the vicinity of the submerged dam and baffle dike are located in Figure 2.5-16. The logs of the borings are presented in Figures 2.5-162 through 2.5-221. Laboratory testing was performed on representative bulk samples of the borrow materials and in situ material. Results of the tests are presented in the following tables and figures: a. optimum moisture contents: Table 2.5-36; b. particle size analyses of placed material: Figure 2.5-424; c. field compaction tests: Figure 2.5-423. d. triaxial compression tests: Tables 2.5-11 and 2.5-17; Figures 2.5-404 through 2.5-406 and 2.5-408; e. unconfined compression tests: Figures 2.5-204, 2.5-205, 2.5-207, 2.5-210 through 2.5-214, and 2.5-216 through 2.5-221; f. dynamic triaxial compression tests: Table 2.5-22; Figures 2.5-411 and 2.5-412; g. resonant column tests: Table 2.5-27; h. Atterberg limit of placed material: Tables 2.5-60 and 2.5-36; and

i. permeability tests: Tables 2.5-31 and 2.5-33. 2.5.5.3.3 Section E-E' along the North Fork of Salt Creek Borings in the vicinity of Section E-E' along the North Fork of Salt Creek are located in Figure 2.5-278. The logs of the borings are presented in Figures 2.5-94 to 2.5-97. Results of laboratory testing of samples are presented in Tables 2.5-6, 2.5-23, and 2.5-43.

CPS/USAR CHAPTER 02 2.5-96 REV. 11, JANUARY 2005 2.5.5.4 Compacted Fill 2.5.5.4.1 Main Dam The specifications for the placement of compacted fill for the main dam are presented in Subsection 2.5.6.4.2.1. 2.5.5.4.2 Ultimate Heat Sink The slopes on the north and south sides of the ultimate heat sink were excavated with no compacted fill material being placed in these areas. The west end of the heat sink is defined by the submerged dam. The compacted fill specifications and construction procedures for the embankment are discussed in Subsection 2.5.6.4.1.1. Compacted fill was also placed at the abutments for the submerged dam and is discussed in Subsection 2.5.6.3.1. A 2-foot thick, 8-inch layered system of soil cement slope protection was provided for the submerged dam, and at the abutments of the submerged dam up to elevation 700 feet. A 3-foot thick layered system of soil element was placed on top of the baffle dike. A typical cross section of the embankment and slope protection is shown in Figure 2.4-24. The east side of the ultimate heat sink is comprised of the screen house and backfill, the baffle dike abutment, the SSWS outlet structure, and the natural excavated slopes. Compacted fill consisting of the Wisconsinan glacial till of the Wedron Formation was used around the screen house and outlet structure as backfill material in the overexcavated areas for these structures.

This operation is discussed in Subsections 2.5.4.5.2.5 and 2.5.4.5.3.5. Compacted fill used for the baffle dike abutment consisted of the same material used to construct the baffle dike as discussed in Subsection 2.5.6.4.1. The natural slopes comprising the remaining portions of the east side were excavated where necessary to provide final slopes of 5:1 (horizontal to vertical) up to elevation 690 and then slopes of 3.5:1 (horizontal to vertical) to natural grade. 2.5.6 Embankments and Dams 2.5.6.1 General 2.5.6.1.1 Ultimate Heat Sink The emergency core cooling system ultimate heat sink was formed by constructing a submerged pond in the valley of the North Fork of Salt Creek. The submerged pond is located immediately west of the screen house and main plant as shown in Figure 2.5-386. The pond's required capacity was developed by constructing an earth dam across the valley bottom and excavating a portion of the valley bottom down to a bed elevation of 668.5 feet. An earthfilled dike was constructed down the center of the pond to lengthen the flow path of the emergency cooling water. The top elevation of the earth dam is 675 feet, and 675 feet for the baffle dike. With the normal operating water level in the lake at elevation 690 feet, the earthfill will be submerged and will not be in use. However, when the water level in the lake drops below elevation 675 feet, the submerged dam will act as an overflow weir and maintain a water level of 675 feet at all times at the intake screen house. The submerged dam was designed to impound the quantity of water necessary for a 30-day supply in order to safely shut down the reactor in the event that the lake level drops below elevation 675 feet.

CPS/USAR CHAPTER 02 2.5-97 REV. 11, JANUARY 2005 2.5.6.1.2 Main Dam The main dam is an earthfill structure constructed to impound the water for the cooling lake. It is located approximately 1200 feet downstream of the confluence of the North Fork of Salt Creek and Salt Creek. The ground surface elevation of the valley bottom at the dam site is approximately 655 feet. The crest of the dam is at 711.8 feet. The normal operating water level in the cooling lake is elevation 690 feet with a probable maximum flood (PMF) level at elevation

708.8 feet. An 80-foot wide concrete service spillway is located on the west side of the dam near the west abutment. The ogee of the service spillway has a crest elevation of 690 feet. In addition to the spillway, there is a low-water intake structure designed to maintain a constant flow downstream of the dam if the lake level falls below elevation 690 and to lower the water level in the lake down to elevation 668 if necessary. On the east side of the main dam, a 1200-foot wide auxiliary spillway was excavated. It was from this area that the material used to construct the dam was obtained. The auxiliary spillway was designed to be operational when the lake level exceeds 700 feet. Figures 2.4-1 and 2.4-13 give a plan view of the dam site illustrating the location of the dam and its appurtenant structures.

2.5.6.2 Exploration 2.5.6.2.1 Ultimate Heat Sink A total of 60 test borings were drilled in the heat sink area as shown in Figure 2.5-16, and described in Subsections 2.5.4.3 and 2.5.5.1.2. The logs of the exploratory borings are shown in Figures 2.5-162 through 2.5-221. The exploration program in the ultimate heat sink showed that the North Fork of Salt Creek has eroded into the Illinoian till of the unaltered Glasford Formation. During glacial and postglacial times, the original stream valley has been filled with alluvial silt and silty clay and outwash sand and gravel. The existing stream flowed on this alluvial and outwash valley fill material. The exploration program indicated that these glaciofluvial deposits ranged in thickness from 12 to 23 feet. The upper 5 to 10 feet consisted of silty clay and clayey silt. The lower part consisted predominantly of medium to coarse sand. Illinoian till underlies the valley fill deposits. The Illinoian till consists of hard, dense, sandy silt and clayey silt and occasional seams and pockets of sand and gravel. The Illinoian till is underlain by pre-Illinoian glacial deposits which are comprised principally of clayey silt and sandy silt. Under the southwest corner of the ultimate heat sink, the pre-Illinoian till is underlain by Mahomet Bedrock Valley deposit. The pre-Illinoian till and the glacial outwash in the southwest part of the ultimate heat sink are underlain by Pennsylvanian bedrock. Pennsylvanian bedrock at the ultimate heat sink consists primarily of shales and siltstones.

Geologic sections are shown in Figure 2.5-284. 2.5.6.2.2 Main Dam A total of 60 borings along the dam alignment and 17 borings in the borrow area were drilled to investigate the subsurface conditions at the dam site. The boring locations are shown in Figure CPS/USAR CHAPTER 02 2.5-98 REV. 11, JANUARY 2005 2.5-272, and a discussion of the drilling operation is presented in Subsection 2.5.4.3. The logs of the exploratory borings are shown in Figures 2.5-74 through 2.5-144. The location of the dam is situated on the floodplain of Salt Creek where glacio-fluvial deposits have filled the original channel cut by the river in the Illinoian till. Wisconsinan through Illinoian age glacial materials form both of the dam abutments. The channel fill was from 16 to 32 feet thick (elevation 627 to 660) where penetrated by borings. Topsoil and alluvial silt and clay ranging from 7 to 18 feet in thickness comprised the upper portion of this material. The lower portion of the alluvial material contained principally outwash sand and gravel interbedded with subordinate zones of silt or clayey silt. The Illinoian till beneath the floodplain deposits, approximately 140 feet in thickness, was composed principally of hard, dense, sandy silt and clayey silt, and occasional intervals of thin layers of sand and gravel. A thin sandy zone in the till was penetrated in Boring D-8 at a depth of 51 feet (elevation 605) or 22 feet below the bottom of the alluvial channel. As presented in Boring D-11, pre-Illinoian glacial deposits, approximately 20 feet in thickness, underlay the Illinoian till. Underlying the pre-Illinoian is a glacio-fluvial sand also about 140 feet thick which comprises the outwash fill of the Mahomet Bedrock Valley. Boring D-11, located 1000 feet downstream from the axis of the main dam, penetrated 50 feet of bedrock consisting of Pennsylvanian limestone, shale, and micaceous siltstone below the valley fill. 2.5.6.2.3 Cooling Lake Reservoir In general, the cooling lake reservoir basin is enclosed by relatively impermeable glacial till. Above elevations ranging from 640 to 700 feet, till of Wisconsinan age and the interglacial zone form the reservoir sidewalls. Below these elevations, Illinoian till prevails and is composed principally of hard, dense, sandy silt and clayey silt, and occasional intervals of permeable sand and gravel a few feet thick. These permeable intervals are not laterally extensive and, therefore, will not provide significant paths for reservoir seepage. The Illinoian till is generally on the order of 150 feet thick. Except for occasional thin and discontinuous lenses of sand, it is low in permeability. The channels of Salt Creek and the North Fork of Salt Creek, where drilled, have incised the Illinoian till to a maximum depth of about 35 feet (elevation 625). Hence, even beneath the stream channels, at least 115 feet of relatively impermeable Illinoian till will separate alluvial channel deposits beneath the reservoir from the deeper Mahomet Bedrock Valley fill of glacial outwash. The highest elevation of the outwash observed in borings is 500 feet. The outwash is underlain by Pennsylvanian bedrock consisting primarily of shales and siltstone from elevation 360 to 400 feet. 2.5.6.3 Foundation and Abutment Treatment 2.5.6.3.1 Ultimate Heat Sink Both the baffle dike and the submerged dam in the ultimate heat sink were founded on sound Illinoian till. The bases of the baffle dike and the submerged dam were founded on the till at approximately elevations 660 feet and 658 feet, respectively. All sand pockets exposed at the surface of the submerged dam excavation were removed. Sand pockets exposed at the surface of the baffle dike excavation were tested and confirmed to have 85% relative density or removed. A comprehensive subgrade testing program under the supervision of the Quality Control personnel at the site was instituted to verify that adequate subgrade was obtained.

CPS/USAR CHAPTER 02 2.5-99 REV. 11, JANUARY 2005 At the location of the north abutment for the submerged dam, the in situ material was excavated to a final slope of 1.5:1 (horizontal to vertical) from the base of the excavation at approximately elevation 660 feet up to elevation 675 feet. Illinoian till was exposed on this excavated face.

Above elevation 675, the in situ material was excavated on a slope of 2:1 (horizontal to vertical) up to existing grade. Following the completion of the excavation, Type A cohesive fill material was placed and keyed into the exposed abutment in near horizontal lifts as described in

Subsection 2.5.6.4.1.1. Likewise, the south abutment for the submerged dam was excavated into in situ material on a slope of 1:2 (horizontal to vertical) to expose Illinoian till between the base of the excavation at approximately 658 feet up to elevation 675 feet. Above this elevation, the in situ material was excavated to a slope of 3:1 (horizontal to vertical) up to existing grade. Type A cohesive fill material was then placed and keyed into the exposed surface in near horizontal lifts. The east abutment for the baffle dike was excavated from the base of the excavation at approximately 660 feet on a slope of approximately 1.5:1 (horizontal to vertical) up to existing grade. A sand lens in the Illinoian till, averaging 6 feet in thickness and ranging up to 14 feet in thickness, was exposed in the abutment. Type A cohesive material was placed and compacted against the sand in near horizontal lifts. Areas adjacent to the abutment where this sand lens was exposed received a blanket of Type A cohesive material to stabilize the slope and control

seepage. No additional foundation or abutment treatment was performed other than that which was described above. The Illinoian till provided a firm, impermeable base for the foundations of both the baffle dike and submerged dam as well as the two abutments for the submerged dam. This, coupled with the relatively impermeable Type A cohesive material used in the embankments, should prevent any appreciable amount of seepage either beneath the baffle dike or under and

around the submerged dam. 2.5.6.3.2 Main Dam The foundation for the main dam consisted of both the Illinoian till and the lower alluvial materials consisting primarily of medium dense to very dense sand with some gravel. Along the centerline of the dam and extending a minimum of 20 feet each side of the centerline, the Illinoian till was exposed and established as the foundation for the key trench. This key trench was excavated to a minimum depth of 2 feet into the Illinoian till to provide a seepage cutoff for the dam. For the wing areas of the dam, the dense sand was used for the foundation. The sand was considered as acceptable subgrade providing it had a minimum relative density of 70%. However, in some areas, the alluvial sands were not suitable as subgrade material and were excavated to expose the sound Illinoian till. A comprehensive subgrade testing program verified that acceptable subgrade was obtained, that Illinoian till was exposed in the keyway, and that the sand had a minimum relative density of 70% as determined by ASTM D-2049.

Figure 2.4-14 shows a typical section of the main dam. The west abutment was comprised of loess, Wisconsinan till, Robein Silt, and weathered and unweathered (unaltered) Illinoian till. These materials have a stiff to very stiff consistency and possess high shear strength. The abutment was excavated on a slope of approximately 1:1 (horizontal to vertical) through these materials down to the Illinoian till of the unaltered Glasford Formation. The service spillway is located adjac ent to the west abutment of the dam and was founded on unweathered Illinoian till. A comprehensive program of subgrade testing was CPS/USAR CHAPTER 02 2.5-100 REV. 11, JANUARY 2005 performed as part of the Quality Control program to verify that acceptable subgrade was obtained. Type A cohesive material was placed and compacted to a minimum dry density of 102% of Standard Proctor on the subgrade beneath the service spillway to bring the foundation up to the design grade. This material was also placed and keyed into the abutment to provide a seepage cutoff. The east abutment was also comprised of loess, Wisconsinan till, Robein silt, and weathered and unweathered (unaltered) Illinoian till. These materials have a stiff to very stiff consistency and high shear strength. The key trench, 40 feet in width, was excavated through the subgrade of alluvial sands exposed in the wing areas of the dam and into the Illinoian till and extended to the abutment. Illinoian till was exposed in the bottom of the key trench. The interglacial materials and the Wisconsinan till were exposed on the face of the abutment slope which was approximately 3:1 (horizontal to vertical). These in situ materials were considered to be suitable subgrade and Type A cohesive material was placed and keyed into them. 2.5.6.4 Embankment 2.5.6.4.1 Ultimate Heat Sink The submerged dam is a homogeneous earth embankment constructed using Type A cohesive fill material. The submerged dam has an average height of 17 feet from the subgrade to the top of the dam at elevation 675 feet. The slopes of the dam are 5:1 (horizontal to vertical). A soil-cement mixture was placed in a layered system over the sides and top of the embankment to a thickness of 2 feet to provide slope protection. A typical section of the submerged dam is illustrated in Figure 2.4-24. The baffle dike is also a homogeneous earth embankment constructed using Type A cohesive material. The baffle dike has an average height of 16 feet from the subgrade to the top of the dike at elevation 676 feet. The side slopes of the dike are 5:1 (horizontal to vertical). The upper 3 feet of the dike embankment is a layered system of soil cement for protection against wave action. A typical section of the baffle dike is illustrated in Figure 2.4-24. 2.5.6.4.1.1 Subgrade and Embankment Materials The foundation material for both the submerged dam and the baffle dike was the Illinoian till of the unaltered Glasford Formation. A comprehensive program for subgrade testing was initiated to verify that the in-place density was greater than 120 pcf. One test was performed for every

10,000 ft 2 of exposed subgrade for the submerged dam and 20,000 ft 2 for the baffle dike. In addition to verifying in-place density, a laboratory sample for classification characteristics was taken. Laboratory testing on subgrade samples taken at each location included a grain-size analysis (ASTM D-422) and Atterberg limit (ASTM D-423 and ASTM D-424). Fill placement for the embankments commenced immediately following subgrade approval. The fill material was classified as Type A cohesive fill material. This material was excavated from the borrow areas outlined in Figure 2.5-384. The material was placed in near horizontal lifts not exceeding 8 inches in loose thickness and compacted by a Hyster C-450 segmented pad sheepsfoot roller. After the completion of each day's work, the top of the embankments, borrow areas, and stockpiles were bladed or rolled smooth and crowned slightly to allow rain to freely run off the surface and prevent ponding.

CPS/USAR CHAPTER 02 2.5-101 REV. 11, JANUARY 2005 A comprehensive laboratory and field testing program was established as an integral part of the Quality Control program at the site. Representative bag samples of the borrow material were

taken for every 6000 yd 3 of material removed to determine the physical characteristics of the material in the laboratory. The following is a list of the tests performed and figures and tables illustrating the average properties of the material used in the embankments: a. grain-size analysis (ASTM D-422) - Figure 2.5-424; b. Atterberg limits (ASTM D-423 and D-424) - Table 2.5-36;

c. Modified Proctor density tests (ASTM D-1557) - Figure 2.5-423; and

d. soil classification (ASTM D-2487 and D-2488). The appropriate specifications for each property are provided on the figures and tables listed above. The Modified Proctor density test was performed once per every 6000 yd 3 of material used.

The field testing program consisted of in-place density and moisture testing using either the Washington densometer method (ASTM D-2167) or the nuclear density method (ASTM D-2922). The in-place density and moisture requirement s were a minimum of 90% of the Modified Proctor maximum dry density and a placed moisture content within 4.5% above the optimum moisture content and 2.0% below the optimum moisture content. The frequency of testing was one test per 10,000 ft 2 per lift. Statistical analyses of the in-place density tests for the Category I cohesive fill used to construct the UHS Dam and Baffle Dike were performed to verify the compaction requirements. The average dry density, average moisture content, and average percent compaction of the fill were determined using the results of the nuclear and sand cone tests performed on the cohesive fill. These averages were computed for each one-foot interval and for the total embankment for both the UHS Dam and the Baffle Dike. A total of 1,526 in-place density tests were analyzed for the UHS Dam. The average dry density of entire UHS Dam embankment was 125.7 PCF placed at an average moisture content of

10.6%. This dry density corresponds to an average of 93.4% compaction as determined by ASTM D1557. Figure 2.5-451 shows the distribution for the dry density test results. The lowest in-place dry density recorded was 119.4 PCF while the highest was 139.4 PCF. Figure 2.5-452 shows the distribution for the in-place moisture content of the tests. The lowest recorded value was 5.7%

and the highest was 14.1%. Figure 2.5-453 shows the distribution of the percent compaction for the in-place tests. The lowest value recorded was 89.1% and the highest value was 103.4%. There were a total of 9 in-place tests that did not meet the 90% degree of compaction requirement. This represents approximately 0.6% of all of the in-place tests taken for the UHS Dam embankment. Their values ranged from 89.1% to 89.9%. Two of the nine tests had both a nuclear test (ASTM D2922) and a balloon test (ASTM D2167) performed at the same location.

In both cases, one of the two tests met the 90% compaction requirement. Two other tests had values of 89.9% compaction. These 9 tests represent isolated areas within the embankment and have in-place compaction densities only slightly less than the minimum requirement. Therefore, they will not be detrimental to the integrity of the UHS Dam embankment.

CPS/USAR CHAPTER 02 2.5-102 REV. 11, JANUARY 2005 There were a total of 5 in-place tests for the submerged dam for which the placement moisture criteria was not met. This represents approximately 0.3% of all the tests. Of these tests, three had moisture contents 0.1% above the specified limit and the other two were 0.2% and 0.6%

above the specified limit. These tests represent only isolated areas and are only slightly outside of the specified limit. Therefore, they were accepted and considered not to be detrimental to the

integrity of the dam. A total of 2034 in-place density tests were analyzed for the Baffle Dike. The average dry density of the entire Baffle Dike embankment was 125.6 PCF placed at an average moisture content of 10.8%. This dry density corresponds to 93.6% compaction as determined by ASTM

D1557. Figure 2.5-454 shows the distribution for the dry density test results for the Baffle Dike. The lowest in-place dry density recorded was 116.6 PCF, while the highest was 143.6 PCF. Figure 2.5-455 shows the distribution for the in-place moisture content of the tests. The lowest recorded value was 6.1% and the highest was 14.8%. Figure 2.5-456 shows the distribution of the percent compaction for the in-place tests. The lowest value recorded was 89.1% and the highest value was 107.6%. There were a total of 7 in-place tests for the Baffle Dike that did not meet the 90% degree of compaction requirement. This represents approximately 0.3% of all of the in-place tests for the Baffle Dike. Their values ranged from 89.1% to 89.9%. Four of the eight tests had a value of 89.9%. Two of the failing tests were taken at the same location as another test (a nuclear test and a baloon test). The second test did meet the 90% compaction requirement in each case.

These 7 tests represent isolated areas within the dike and have in-place compaction densities only slightly less than the minimum requirement. Therefore, they will not be detrimental to the integrity of the Baffle Dike. There were a total of 11 in-place tests for the Baffle Dike which did not meet the placement moisture criteria. This represents approximately 0.5% of all the tests. Of these tests, nine had in-place moisture content that were only 0.1% above the specified limit. One had a moisture content 0.2% above the limit and the other was 0.5% above the limit. The test that was 0.5% above the limit also had another test at the same location and it was within the specified limits.

These tests represent only isolated areas and are not considered to be detrimental to the

integrity of the Baffle Dike. Due to the highly preconsolidated nature of the underlying clays, such as the Illinoian till and the small additional load due to the embankment itself, no measurable consolidation or settlement is anticipated in the subgrade. No appreciable settlement is expected in the embankments themselves due to the small heights of the embankments. 2.5.6.4.1.2 Slope Protection A soil cement mixture was placed on the baffle dike and the submerged dam for slope protection. For the baffle dike, the fill embankment was built to elevation 673 feet. The remaining 3 feet of the dike was constructed using the soil cement. It was placed in maximum 8-inch loose lifts and compacted using a smooth wheel (drum) compactor. The compacted material was then tested to verify that the specifications were met. Each succeeding lift was smaller in width to provide a stair-step effect to dissipate any wave action.

CPS/USAR CHAPTER 02 2.5-103 REV. 11, JANUARY 2005 For the submerged dam, the soil cement was placed in maximum 8 inch loose lifts and compacted to a total 2 foot thick layer. It was placed parallel to the slopes which provided a relatively smooth surface for the outside of the embankment. Immediately downstream of the submerged dam, an area 56 feet wide at approximately elevation 670 feet received a 2-foot layered system of soil cement to prevent the possibility of undermining the dam. Also, the soil cement slope protection was placed and compacted on the north and south abutments of the submerged dam to protect them from scouring. No slope protection was provided for the baffle dike abutment which had a final slope of 4:1 (horizontal to vertical) above elevation 676 feet. The soil cement was a mixture of soil aggregate, cement and water mixed to proper proportions at a central batching plant. The material was then transported and placed using dump trucks. The material was spread on a moist surface by either a pavement spreading machine or bulldozer. The in-place density requirements were a minimum dry density of 95% of the ASTM D-558 method with a moisture content within 1% above the optimum moisture content and 2%

below the optimum. The test frequency was one test per 10,000 ft 2 per lift. Statistical analyses of the in-place density tests for the soil cement slope protection for the UHS Dam and Baffle Dike were performed to verify the compaction requirements. The average dry density, the average moisture content, and the average percent compaction of the soil cement were determined using the results of the nuclear and sand cone tests performed. These averages were computed for the 1.) Baffle Dike, 2.) Slopes and crest of the Submerged Dam, and 3.) Submerged Dam including the abutments and downstream flat area. A total of 126 in-place density tests were analyzed for the Baffle Dike. Figure 2.5-457 shows the distribution for the dry density test results. The average dry density was 130.1 PCF with the lowest value being 125.8 PCF and the highest value being 137.9 PCF. Figure 2.5-458 shows the distribution for the in-place moisture content. The average moisture content was 9.4% with the lowest value being 7.5% and the highest value being 11.1%. Figure 2.5-459 shows the distribution for the percent compaction for the in-place tests. The average percent compaction was 98.1% with the lowest value being 94.3% and the highest 104.4%. There were 2 tests that did not meet the 95% compaction requirement, which were averaged with 2 other adjacent in-place tests. The average of each of these tests was greater than the 95% compaction requirement. A total of 119 tests were performed on the crest, upstream slope, and downstream slope of the UHS Submerged Dam. Figure 2.5-460 shows the distribution for the results of the dry density tests. The average dry density was 126.8 PCF with the lowest value being 122.8 PCF and the highest value being 131.0 PCF. Figure 2.5-461 shows the distribution for the moisture content test results. The average moisture content was 10.7% with the lowest value being 9.3% and the highest value being 12.3%. Figure 2.5-462 shows the distribution for the percent compaction for the tests. The average percent compaction was 97.0% with the lowest value being 94.9% and the highest value being 99.9%. There was only 1 in-place test that did not meet the 95% compaction requirement. This test was averaged with two additional adjacent tests. This average was greater than 95% compaction and was considered acceptable. A total of 205 in-place density tests were analyzed when the tests performed on the dam abutments and downstream area were included in the summary. Figure 2.5-463 shows the distribution for the dry density test results. The average dry density was 126.7 PCF with the CPS/USAR CHAPTER 02 2.5-104 REV. 11, JANUARY 2005 lowest value being 102.4 PCF and the highest being 132.4 PCF. Figure 2.5-464 shows the distribution for the moisture content test results. The average moisture content was 10.6% with the lowest value being 8.7% and the highest value being 12.3%. Figure 2.5-465 shows the distribution for the percent compaction for the test results. The average percent compaction was 96.5% with the lowest value being 77.5% and the highest value being 100.2%. A total of 11 in-place tests did not meet the 95% compaction requirement. Three of these tests were averaged with two additional adjacent tests. The averages of these tests were greater than 95% compaction and were considered acceptable. The remaining 8 failing tests were performed on the downstream flat area immediately downstream of the submerged dam. These tests were taken on the middle layer of three placed in this area and were documented in a non-conformance report. This area was considered acceptable based on the fact that the layer is confined between two acceptable layers, appeared to be hard and solid on the surface, and would not be detrimental to the integrity of the submerged dam. A total of 4 in-place tests performed on the soil cement did not meet the specified moisture content limits. All four were in the Baffle Dike area. Three of these were in the same area (one initial failure and two retests). The average moisture content of these three tests was 0.3% below the specified limits. The one other failing te st value was averaged with two retests. The resulting average of these was within the specified limits. The percent compaction for all of these tests was acceptable. These failing tests represent only minor areas and they will not be detrimental to the integrity of the Baffle Dike. Representative bag samples of the soil aggregate for the soil cement were obtained from the stockpile for every 4000 yd 3 of material used. A grain-size analysis (ASTM D-422) was performed on each sample to verify that the material met the specifications. Also, a five-point compaction test (ASTM D-588) was performed on representative samples for every 20,000 yd 3 of material placed. This test was used to develop a family of curves from which maximum dry densities and optimum moisture contents could be determined. A three-point compaction method was used for the field control of in-place density and moisture requirements. A representative sample of placed soil cement was obtained and three compaction tests were performed at three different moisture contents: one as placed, one above, and one below. The curve defined by these three points was compared to the family of curves and a maximum dry density and optimum moisture content was then determined. This test procedure was performed for every 4000 yd 3 of material placed. 2.5.6.4.2 Main Dam The main dam is a homogeneous earth embankment constructed using Type A cohesive fill material. The top of the dam is at elevat ion 711.8 feet with the average ground elevation at 655 feet, thus giving an average effective height of t he dam of approximately 56.8 feet. The dam is approximately 3000 feet in length, with side slopes of 3:1 (horizontal to vertical). Riprap material was placed as slope protection on portions of both the upstream and downstream faces of the dam. A typical section of the main dam is presented in Figure 2.4-14. 2.5.6.4.2.1 Subgrade and Embankment Materials The foundation materials for the main dam consisted of the Illinoian till and the dense alluvial sands. A comprehensive subgrade testing program for the main dam was initiated as described in Subsection 2.5.6.3.2. One in-place density test was performed for every 20,000 ft 2 of CPS/USAR CHAPTER 02 2.5-105 REV. 11, JANUARY 2005 exposed subgrade. Laboratory testing on subgrade samples taken at each test location included: grain-size analysis (ASTM D-422) and Atterberg limit (ASTM D-423 and D-424) on cohesive samples, and relative density (ASTM D-2049) for granular samples. Fill placement commenced immediately following subgrade approval. The borrow material was obtained from the excavation for the auxiliary spillway on the east side of the dam. The material used in the embankment was placed in near horizontal lifts not exceeding 8 inches in loose thickness and was compacted by heavy sheepsfoot roller compaction equipment. Following the completion of each day's work, the top of the embankment and borrow areas were bladed smooth and crowned slightly to promote surface runoff in the event of rain. A 2-foot thick blanket drain was constructed in the embankment on the downstream side of the dam to control seepage. This is discussed in Subsection 2.5.6.6.2. The subgrade for the service spillway was established on the Illinoian till at approximately elevation 667 feet and at elevation 630 feet for the stilling basin. The subgrade testing program previously described for the dam was also followed for this area. Due to the overexcavation to obtain suitable subgrade on the upstream side of the service spillway, a maximum of 12 feet of Type A fill was placed and compacted to a minimum dry density of 102% of the Standard Proctor maximum dry density as determined by ASTM D-698. This material was placed in near horizontal lifts in loose thicknesses not exceeding 6 inches and was compacted using heavy sheepsfoot roller compaction equipment. The placement moisture content was between the optimum water content and 4% below the optimum. A comprehensive laboratory and field testing program was established as an integral part of the Quality Control program at the site. Representative bag samples of the borrow material were taken for every 10,000 yd 3 of material placed to determine the physical characteristics of the material in the laboratory. Subsection 2.5.5.3.1 describes both the preliminary testing and field testing in conjunction with the construction of the dam. The field testing program consisted of in-place density and moisture testing using either the Washington densometer method (ASTM D-2167) or the nuclear density method (ASTM D-2922). The in-place density and moisture requir ements were a minimum of 95% of the maximum dry density determined by ASTM D-698 and a placed moisture content within 2.0%

above the optimum moisture content and 4.0% below the optimum moisture content. The frequency of testing was one in-place density test per 10,000 ft 2 per lift. Due to the highly preconsolidated nature of the underlying soils, such as the Illinoian till, no measurable consolidation or settlement is anticipated in the subgrade areas where the Illinoian till was exposed. In the areas where the dense alluvial sands formed the subgrade, any consolidation or settlement in these materials was expected to occur immediately and simultaneously with the construction of the embankment. 2.5.6.4.2.2 Slope Protection Riprap material was placed on both the upstream and downstream faces of the main dam. On the upstream face of the dam, 18 inches of riprap was placed over two 9-inch layers of graded bedding materials. The riprap and bedding materials were placed from elevation 677 feet up to elevation 704 feet. From elevation 704 feet up to the top of the dam, only the riprap and one CPS/USAR CHAPTER 02 2.5-106 REV. 11, JANUARY 2005 18-inch to 9-inch bedding layer was placed. This design allowed for 13 feet of placed riprap beneath the normal pool elevation of 690 feet on the upstream face. On the downstream face of the dam, riprap was placed from the toe of the dam up to elevation 670 feet. An 18-inch thick layer of the riprap was placed over two 9-inch layers of graded bedding material. Above elevation 670, a 4-inch layer of topsoil was placed and seeded for erosion control. The riprap and related filters were designed following the recommendations of the U.S. Bureau of Reclamation (Reference 88) and the U.S. Army Corps of Engineers (Reference 89). The maximum size of the riprap was 30 inches while the average size was approximately 9 inches. 2.5.6.5 Slope Stability 2.5.6.5.1 Ultimate Heat Sink The static and dynamic slope stability analyses for the submerged earthfill and baffle dike are discussed in Subsections 2.5.5.2.2, 2.5.5.2.3, and 2.5.5.2.4. The results of the laboratory testing are described in Subsection 2.5.5.3.2. 2.5.6.5.2 Main Dam The slope stability analysis for the main dam is presented in Subsection 2.5.5.2.1. The results of the laboratory testing are presented in Subsection 2.5.5.3.1. The updated parameters used in the stability analysis were determined from a program of in-place testing involving the extraction of shelby tube samples from the embankment in the spring of 1977. These values were used to verify the design parameters determined by the initial laboratory tests. Triaxial testing was performed on these samples to determine the actual in situ strength obtained during construction. This program is discussed in Subsection 2.5.6.9. 2.5.6.6 Seepage Control 2.5.6.6.1 Ultimate Heat Sink Field permeability tests were performed in the Illinoian till with the results being presented in Table 2.5-38. Also, laboratory permeability tests were performed on samples of the Illinoian till and are presented in Table 2.5-31. These tests indicate that the highest permeability for the subgrade materials for the submerged dam is 1.4 x 10

-5 cm/sec. Laboratory permeability tests were performed on remolded samples of the fill material. A coefficient of permeability of 2 x 10

-8 cm/sec was calculated for the embankment material. The dam and baffle dike were designed as submerged structures, to be operative only in the event that the cooling lake is lost. When in operation, the dam will only have a head of approximately 7 feet and seepage is expected to be minimal during the shutdown period. Due to the combination of low permeabilities, small head, and long seepage path, no significant amount of seepage is expected from the ultimate heat sink during the 30 day shutdown period when the heat sink is operational.

CPS/USAR CHAPTER 02 2.5-107 REV. 11, JANUARY 2005 2.5.6.6.2 Main Dam Laboratory and field permeability testing was performed on representative samples of in situ and remolded materials for the surrounding vicinity and the dam embankment. Coefficients of permeabilities for in situ material are presented in Tables 2.5-32 and 2.5-38 for the laboratory and field tests, respectively. The coefficients of permeability for remolded samples are presented in Table 2.5-33. The estimated coefficient of permeability for the embankment

materials was 2 x 10

-8 cm/sec. A cutoff trench was constructed along the centerline of the dam to control the seepage beneath the dam. Because most of the dam was constructed on dense sands, the cutoff trench was excavated a minimum of 2 feet into the Illinoian till and had a minimum width of 40 feet. The cutoff trench was then filled with Type A fill material compacted in accordance with the specifications for the embankment fill. As shown in Figure 2.4-14, a 2-foot thick sand blanket drain was placed from the downstream toe of the dam into the dam for a distance of 100 feet. This drain will prevent any seepage from emerging from above the immediate toe of the dam. A collector ditch was established approximately 20 feet from the toe of the dam to collect any seepage from the dam and control its flow away from the dam. 2.5.6.7 Diversion and Closure 2.5.6.7.1 Ultimate Heat Sink The North Fork of Salt Creek was diverted through a channel excavated along the north side of the heat sink and across the dam centerline to allow the excavation and fill placement for the baffle dike and the submerged dam. A slurry trench was installed and a temporary dike was constructed along the south side of the diversion channel and around the west side of the submerged dam to prevent seepage into the excavations. For the section of the submerged dam north of the diversion channel, a slurry trench was constructed in a U-shape with the two open ends keying into the abutment. Seepage into the excavations was collected in a series of sumps and was pumped into the diversion channel as it became necessary. After the baffle dike was completed and the north and south sections of the submerged dam were approximately 95% complete, a new diversion channel was excavated through the newly placed fill material in the north section of the dam. Earth cofferdams were constructed

immediately upstream and downstream of the submerged dam where the original diversion channel flowed across the centerline of the dam. The new diversion channel allowed the flow in the North Fork to continue while the center section of the dam could be dewatered by pumping, excavated to suitable subgrade, and fill placed to tie in the two existing fill sections of the dam. After the completion of this section and the placement of the soil cement slope protection on the baffle dike and dam, a temporary cofferdam was constructed across the new diversion channel for the final closure of the dam and the impoundment of water in the heat sink. Natural ground in this area was above elevation 675 feet, being outside the limits of the heat sink, thus allowing this area to remain relatively dry while the heat sink was filling. Debris and overly wet material were removed from the new diversion channel prior to fill placement.

CPS/USAR CHAPTER 02 2.5-108 REV. 11, JANUARY 2005 2.5.6.7.2 Main Dam A slurry trench was used to prevent seepage from entering into the excavations for the main dam. The trench was U-shaped for each section of the dam on either side of Salt Creek and the open ends keyed into the abutments. The excavation for the cutoff trench along the dam centerline was performed first to allow the alluvial sands to drain to permit the excavation to the suitable subgrade sand in the dry. The seepage was collected in a sump from where it was pumped back into Salt Creek. After both sections of the embankment were above approximately elevation 670, two earth cofferdams were constructed immediately upstream and downstream of the dam section in the original creek bed. The creek flow was then diverted over the existing fill in a channel excavated approximately 300 feet west of the original creek channel. This final section of the dam was dewatered by pumping, excavated to suitable subgrade, and fill was placed to tie into each of the other embankment sections. The fill was placed only up to approximately elevation 655 feet. At this time, the creek was diverted back to its original course over the newly placed fill in a new channel. When both sections of the embankment were completed to approximately elevation 700 feet, the creek flow was stopped and fill placement of the closure section began. Debris and overly wet material was removed from the closure section prior to fill placement.

2.5.6.8 Performance Monitoring 2.5.6.8.1 Ultimate Heat Sink Monitoring Program The ultimate heat sink (UHS) monitoring program consists of a visual inspection of the UHS shoreline (including the abutments of the UHS submerged dam) to detect scour or erosion around the UHS, a sediment survey of the UHS and immediate area upstream, a hydrographic survey of the UHS submerged dam, and a physical inspection of the submerged dam, the shutdown service water (SSW) outlet structure, and intake screenhouse. The sedimentation

monitoring program for the ultimate heat sink is also discussed in Subsection 2.4.11.6. The monitoring program shall be performed on an annual basis, except for sedimentation accumulation monitoring which shall be done periodically as required to effectively maintain the required UHS Volume. Sedimentation monitoring shall be scheduled based upon as-left results of dredging, past experience with accumulation rates, recent trends, and latest projections.

Annual monitoring shall recommence when approximately 60% to 70% of allowable silt by volume is accumulated, or if sedimentation is approaching elevation 672 feet (msl) at more than two adjacent sections. Additional inspections shall be performed if a major flood, drought, or earthquake occurs. A major flood shall be defined as the 100-year or greater flood. A major drought shall be defined as the 100-year or greater drought for which the lake level is at elevation 682.3 feet. A major earthquake shall be defined as the operating basis earthquake (OBE) having a horizontal peak ground acceleration of 0.10g or larger at the site. 2.5.6.8.1.1 Monitoring Requirements 2.5.6.8.1.1.1 Inspection of UHS Shoreline The UHS shoreline shall be observed from both the land and water (by boat) to determine if scour, erosion, or slope instability has occurred. Scour and erosion shall be defined as any CPS/USAR CHAPTER 02 2.5-109 REV. 11, JANUARY 2005 area where the surface vegetation has been disturbed in an area greater than 100 square feet. Slope instability shall be defined as any area where cracks or dislodged trees are noticeable for a distance of 10 feet along the slope. When applicable, photographs and detailed sketches of any questionable areas shall be documented in the inspection report for comparison of conditions from past and future inspections. The abutments of the UHS submerged dam at and above the water line shall be inspected from a boat for scour and erosion beneath and around the soil cement slope protection of the dam.

Photographs and detailed sketches, when made, shall be documented in the inspection report. 2.5.6.8.1.1.2 Monitoring of UHS Submerged Dam A permanent horizontal and vertical control system will be established on shore as shown on Figure 2.5-485 to provide a fixed guidance for the dam survey.

Annually, a fathometer will be used to measure and record continuous elevations along the crest of the submerged dam. Continuous elevation readings will also be taken for transverse sections of the submerged dam at intervals of 250 feet, measured from the permanent control system, alternating annually between 225 feet, 475 feet, 725 feet, etc., and 100 feet, 350 feet, 600 feet, etc. (Figure 2.5-485). Readings for transverse sections at Stations 3+50, 11+00, and 18+50 will be taken during each survey for a continuous record and comparison of the UHS dam slopes. Whenever the fathometer readings indicate that an area of the crest of the submerged dam is 0.5 feet below the elevation of the crest as established in the 1982 survey, a resurvey will be performed. If the results of the resurvey concur with the first survey, an evaluation of the capacity of the UHS will be made to ensure that the UHS still has its 30-day minimum capacity. The minimum UHS volume, and the plant response to insufficient UHS volume, is given in the Technical Specifications. All fathometer readings will be taken within 5 feet of the designated area and will be accurate to

+/-0.1 feet. Horizontal and vertical control for the annual monitoring of the movement of the UHS submerged dam is accomplished in accordance with Figure 2.5-485 and with the use of a transit, a level, an EDM (Electronic Distance Measuring), a boat equipped with a recording fathometer. Lines of traverse for the boat are established and marked by buoys in the water.

Control is maintained from preset stations on shore for the centerline of the submerged dam and the appropriate cross sections. With the use of the EMD and the fathometer, a continuous strip chart recording the elevation of the top of the dam at centerline is produced to give a profile for the entire dam. Cross sections taken at 250-foot intervals in a similar manner are developed by the fathometer. From the data collected to date, it is evident that no measurable movement has taken place. Due to the submerged dams very flat cross section (1 vertical to 5 horizontal slopes), its 2-foot thick soil-cement facing and its load equilibrium, no movement is expected. Based on the

above, it is known that the submerged dam is intact and is capable of performing its safety function. The UHS dam is to be monitored for movement annually as described above except that cross sections will be taken between those originally taken, every other year, to more completely CPS/USAR CHAPTER 02 2.5-110 REV. 11, JANUARY 2005 envelop the structure. Three cross sections, Station 3 + 50, 11 + 00' and 18 + 50, will be monitored every year to provide another form of comparison for movement. Figure 2.5-485 shows all of the cross sections. All of the abov e survey will be performed by licensed surveyors or licensed professional engineers experienced in this type of work (Q&R 241.22). 2.5.6.8.1.1.3 Sedimentation Monitoring of the UHS and Upstream Area The initial survey of the UHS was performed in 1977 by using aerial photography. A topographic map at a scale of 1 inch equal to 100 feet with 2-foot contour intervals was prepared. From this map the initial volume of the UHS was computed. Permanent horizontal and vertical control points have been established around the perimeter of the UHS to provide the base for the sedimentation monitoring program for the UHS. The 30 designated monitoring locations define nine North/South cross sections of the UHS as shown on Figure 2.4-31 which will be located from these control points. A Raytheon Model DE719B survey fathometer, or equivalent, will be used to measure the bottom elevations of the UHS at these designated cross sections. The amount of sedimentation will be calculated based on the results of these measurements. All fathometer readings will be taken with 5 feet of the designated locations and will be accurate to

+/-0.1 feet. The sedimentation monitoring program will be performed periodically as described in Subsection 2.5.6.8.1. Whenever it is determined that the volume of sediment deposit in the UHS has reached 218 acre-feet, a dredging program will be undertaken. 2.5.6.8.1.1.4 Visual Inspection of Underwater Structures A diver will inspect the concrete structures (SSW outlet structure and the intake screenhouse) for silt or debris accumulation at the intake or discharge points. The concrete structures will also be inspected for concrete deterioration, structural cracking, foundation undermining resulting from scour, and movement along construction joints. Whenever structural cracking, concrete deterioration of more than 0.5 inches in depth over an area larger than 5 square feet is noticed in any UHS concrete structure, or scour beneath a UHS structure of more than 1 foot for a distance of 5 feet is reported, an evaluation of its effect on the structure will be made and documented. If it is determined that the condition is detrimental to the safe operation of the UHS, thus affecting the shutdown service water system, the plant may go into safe shutdown as directed by the Technical Specifications until the situation is remedied. Photographs or detailed sketches of significant conditions will be made to allow an evaluation by qualified engineers. 2.5.6.8.1.1.5 Qualification of Inspection and Evaluation Personnel The inspection and evaluation work will be performed under the direction of qualified engineers by personnel qualified for the particular inspections. They shall be able to identify signs of distress (slope instability, erosion, and deterioration of slope protection and concrete structures) and provide recommendations for corrective actions.

CPS/USAR CHAPTER 02 2.5-111 REV. 11, JANUARY 2005 2.5.6.8.1.2 Technical Evaluation The results of the monitoring programs will be summarized in a report. The results will be compared with the initial and previous reported conditions. Abnormal hazardous conditions observed during the inspection shall be reported immediately to the consulting engineers and to the NRC. The minimum UHS volume, and the plant response to insufficient UHS volume, is given in the Technical Specifications.

2.5.6.8.1.3 Content of Inspection Report A technical report shall be prepared to present the results of each inspection. The original report shall be kept onsite for the life of the plant. The inspection report shall include the following: 1. Statements as to the reason for the survey, such as annual report, major flood, earthquake, etc. 2. Results of visual inspections. 3. Results of the sedimentation monitoring survey.

4. Results of the UHS submerged dam and SSW outlet and intake structures monitoring program. 5. Comparisons at recently performed profiles to the initial and previous profiles of the UHS. 6. Assessments of the causes of abnormal conditions, evaluation of the UHS integrity, and recommendations for additional investigations, remedial measures, or future inspections, where appropriate. 7. A log of abnormal conditions (floods, seismic events, extreme reservoir drawdown, etc...) since the previous inspection. 8. Name and license numbers of the surveyors, engineers, and other personnel that performed the field work and report input. 9. Description of the boat, equipment, lake and weather conditions during the surveys. 10. Summary of daily lake level readings recorded by the owner. 2.5.6.8.1.4 Quality Assurance Requirements The quality assurance/control procedures, special process procedures, and documentation that apply to the work will outline the testing, inspection, etc., activities necessary for the accomplishment of the work and the assurance of its quality and requisite documentation.

CPS/USAR CHAPTER 02 2.5-112 REV. 11, JANUARY 2005 2.5.6.8.2 Main Dam An instrumentation plan was established for the main dam to monitor seepage, settlement, and horizontal movement. Three sets of six hydrostatic pressure cells were installed in the dam to monitor the hydrostatic pressure at various elevations in the dam. Twelve open standpipe piezometers were installed downstream of the dam to depths of 50 feet to monitor groundwater levels in the Illinoian till beneath the dam and in the alluvium and fill adjacent to the dam (Figure 2.5-272 and Table 2.4-31). Slightly elevated hydrostatic pressures were observed in piezometers in the Illinoian till near the left abutment (Figures 2.4-48 through 2.4-50). A system of five relief wells (three presently working) were installed in October 1979 to reduce the pressures. Seven additional piezometers were installed near these relief wells to monitor their performance. Additional discussion of the observation wells is presented in Subsections 2.4.13.2 and 2.4.13.4. Three cross-arm reference points were established along the dam centerline to monitor the settlement in the embankment itself. Also, a system of triangulation points and surface monuments were established to monitor any horizontal or vertical movement of the embankment. 2.5.6.9 Construction Notes An independent testing program was established in the spring of 1977 to determine the depth and extent of frost penetration in the completed portions of the heat sink and main dam embankments. Shelby tube samples were taken at depths up to 4 feet. Testing for these samples included: Atterberg limit, dry density, moisture content, and consolidated-undrained triaxial compression tests. Conservative values for the effective strength parameters used in the slope stability analyses were determined from this testing as shown in Table 2.5-52. From this testing program it was determined that below a depth of 18 inches, the properties of the compacted fill met the design requirements. Therefore, the upper 12 inches were removed and the remaining 6 inches were disced open and recompacted with the heavy sheepsfoot compaction equipment before fill operations resumed in 1977. 2.5.6.9.1 Ultimate Heat Sink The overexcavation for the submerged dam on the downstream side was backfilled with Type C material up to approximately elevation 668.5 feet. This material was obtained from the same borrow areas as the Type A material used in the embankments and it met the specification

requirements of the Type A material. It was placed at moisture contents between 6% above the optimum moisture content and 3% below the optimum moisture content and compacted to a minimum dry density of 85% of the maximum dr y density as determined by the modified Proctor compaction test (ASTM D-1557). For a distance of 56 feet downstream of the toe of the submerged dam, 2 feet of soil cement slope protection was placed over this Type C material. This soil cement was placed to protect the toe of the dam from being undermined in the event the main dam was lost. Beyond the 56-foot distance, random fill material was placed up to existing grade at approximately elevation 673.5 feet. A typical section of the submerged dam is shown in Figure 2.4-24. Soil cement slope protection was used instead of riprap in the heat sink area to protect the surfaces of baffle dike and the submerged dam. The soil cement material proved to be more CPS/USAR CHAPTER 02 2.5-113 REV. 11, JANUARY 2005 economically feasible as well as having sources of suitable borrow material more accessible. Soil cement is discussed in Subsection 2.5.6.4.1.2. The final locations of the north abutment for the submerged dam and the abutment for the baffle dike were changed from their preliminary locations. Additional borings as discussed in Subsection 2.5.4.3 were taken to establish more suitable soil conditions. The alignment of the submerged dam was changed for the northern section of the dam. This realignment permitted the dam abutment to be keyed into unweathered Illinoian till up to elevation 675 feet. The new alignment also prevented encroachment of the abutment fill into the right-of-way for Illinois Highway 54. This realignment is shown in Figure 2.5-384. Geologic mapping for the abutment is discussed in the mapping report presented in Attachment C2.5. Type A cohesive material was not obtained from the station excavation as originally planned. Borrow areas south of the heat sink were determined to be able to provide a more homogeneous and greater quantity of suitable borrow material. These areas are discussed in

Subsection 2.5.6.4.1.1. 2.5.6.9.2 Main Dam Although the major portion of the main dam foundation was established on the dense alluvial sand, it became necessary to remove some areas of the sand that were too loose to be disturbed during construction. In these areas, the sand was completely removed and the Illinoian till was established as the subgrade. Unsuitable borrow material was spoiled in designated areas both upstream and downstream of the main dam. Also, waste material was spoiled adjacent to the upstream slope of the dam up to elevation 670 feet. This material adds more stability to the upstream slope of the dam as having the effect of flattening the final slope to more than 3:1 (horizontal to vertical). 2.5.6.10 Operational Notes The main dam was closed and impoundment of water in Lake Clinton commenced on October 12, 1977. The heat sink dam was closed on October 15, 1977, and had water flowing over the crest on October 21, 1977. No adverse effects of sloughing of the slopes at either dam location

have been recorded. 2.5.7 References for Section 2.5

1. P. B. King, "The Evolution of North America," 190 pp., Princeton University Press, Princeton, New Jersey, 1959. 2. H. B. Willman, et al., "Handbook of Illinois Stratigraphy," Bulletin 95, 261 pp., Illinois State Geological Survey, Urbana, Illinois, 1975. 3. H. Faul, "Ages of Rocks, Planets, and Stars," 109 pp., McGraw-Hill, New York, 1966.

4. J. Boellstorf, "Proposed Abandonment of pre-Illinoian Pleistocene Stage Names," Geol. Soc. America, North Central Section 12th Annual Meeting, Abstracts, p. 247, 1978.

CPS/USAR CHAPTER 02 2.5-114 REV. 11, JANUARY 2005 5. H. B. Willman, and J. C. Frye, "Pleistocene Stratigraphy of Illinois," Bulletin 94, 204 pp., Illinois State Geological Survey, Urbana, Illinois, 1970. 6. K. Piskin and R. E. Bergstrom, "Glacial Drift in Illinois: Thickness and Character," Circular 490, 35 pp., Illinois State Geological Survey, Urbana, Illinois, 1975. 7. J. C. Frye, et al., "Glacial Tills of Northwestern Illinois," Circular 437, 45 pp., Illinois State Geological Survey, Urbana, Illinois, 1969. 8. G. S. Boulton, "Modern Arctic Glaciers As Depositional Models For Former Ice Sheets,"

Quarterly Journal of the Geological Society of London, Vol. 128, pp. 361-393, London, England, 1972. 9. W. B. Howe and J. W. Koenig, "The Stratigraphic Succession in Missouri," Second Series, Vol. XL, 185 pp., Missouri Geological Survey and Water Resources, Rolla, Missouri, 1961. 10. R. H. Shaver, et al., "Compendium of Rock-Unit Stratigraphy in Indiana." Bulletin 43., 229 pp., Indiana Department of Natural Resources, Geological Survey, Bloomington, Indiana, 1970. 11. H. L. James, Stratigraphic Commission Note 40, Subdivision of the Precambrian, an interim scheme to be used by the U.S. Geological Survey, Bull. Amer. Assoc. Petrol. Geologists, V. 56, pp. 1128-1133, 1972. 12. L. N. Stout and D. Hoffman, "An Introduction to Missouri's Geologic Environment," Educational Series No. 3, 40 pp., Missouri Geological Survey and Water Resources, Rolla, Missouri, 1973. 13. H. R. Schwalb, "Paleozoic Geology of the Jackson Purchase Region, Kentucky," Report of Investigations 10, 40 pp., Kentucky Geological Survey, Lexington, Kentucky, 1969. 14. E. Atherton, "Tectonic Development of the Eastern Interior Region of the United States,"

Background

Materials for Symposium on Future Petroleum Potential of NPC Region 9, Illinois Petroleum 96, pp. 29-43, Illinois State Geological Survey, Urbana, Illinois, 1971. 15. A. J. Eardley, "Structural Geology of North America," 624 pp., Harper & Brothers, New York, 1951. 16. J. W. Koenig, "The Lincoln Fold in Northeastern Missouri," Report of Investigations No. 27, pp. 75-80, Missouri Geological Survey and Water Resources, Rolla, Missouri,

September 1961. 17. M. H. McCracken, "Structural Features of Missouri," Report of Investigations No. 49, 99 pp., Missouri Geological Survey and Water Resources, Rolla, Missouri, 1971. 18. K. E. Clegg, "The La Salle Anticlinal Belt in Illinois," Guidebook Series No. 8, pp. 106-110, Illinois State Geological Survey, Urbana, Illinois, 1970.

CPS/USAR CHAPTER 02 2.5-115 REV. 11, JANUARY 2005 19. D. H. Swann and A. H. Bell, "Habitat of Oil in the Illinois Basin," Bulletin of the American Association of Petroleum Geologists, Vol. 42, pp. 447-472, American Association of Petroleum Geologists, Tulsa, Oklahoma, 1958. 20. A. H. Bell, et al., "Deep Oil Possibilities of the Illinois Basin," Circular 368, 38 pp., Illinois State Geological Survey, Urbana, Illinois, 1964. 21. H. B. Willman and J. N. Payne, "Geology and Mineral Resources of the Marseilles, Ottawa, and Streator Quadrangles," Bulletin 66, 388 pp., Illinois State Geological Survey, Urbana, Illinois, 1942. 22. T. C. Buschbach, Illinois State Geological Survey, written communication, 1973. Unpublished Notes, Attachment D2.5, CPS-FSAR. 23. H. S. McQueen, N. S. Hinchey, and K. Aid, "The Lincoln Fold in Lincoln, Pike, and Ralls Counties, Northeastern Missouri," Report of Investigations No. 27, Missouri Geological Survey and Water Resources, Rolla, Missouri, pp. 81-85, 1961. 24. A. H. Bell, "Subsurface Structure of the Base of the Kinderhook-New Albany Shale in Central and Southern Illinois," Report of Investigations 92, 13 pp., Illinois State Geological Survey, Urbana, Illinois, 1943. 25. E. P. DuBois and R. Siever, "Structure of the Shoal Creek Limestone and Herrin (No. 6) Coal in Wayne County, Illinois," Report of Investigations 182, 7 pp., Illinois State Geological Survey, Urbana, Illinois, 1955. 26. W. W. Rubey, "Geology and Mineral Resources of the Hardin and Brussel Quadrangles," Professional Paper 218, 179 pp., U.S. Geological Survey, U.S.

Government Printing Office, Washington, D.C., 1952. 27. W. L. Calvert, "Sub-Trenton Structure of Ohio, with Views, on Isopach Maps and Stratigraphic Sections as Basis for Structural Myths in Ohio, Illinois, New York, Pennsylvania, West Virginia, and Michigan," American Association of Petroleum Geologists Bulletin, Vol. 58, No. 6, pp. 957-972, American Association of Petroleum Geologists, Tulsa, Oklahoma, 1974. 28. D. R. Kolata and T. C. Buschbach, "Plum River Fault Zone of Northwestern Illinois," Circular 491, 20 pp., Illinois State Geological Survey, Urbana, Illinois, 1976. 29. G. A. Ludvigson et al., "A Field Guide to the Plum River Fault Zone in East-Central Iowa," in R. R. Anderson, editor, 42nd Annual Tri-State Geol. Field Conf. Guidebook, Geology of East-Central Iowa, pp. I I-49, 1978. 30. D. R. Kolata et al., "The Sandwich Fault Zone of Northern Illinois," Illinois State Geological Survey, Circular 505, 1978. 31. A. V. Heyl, et al., "The Geology of the Upper Mississippi Valley Zinc-Lead District," Professional Paper 309, 310 pp., U.S. Geological Survey, U.S. Government Printing

Office, Washington, D.C., 1959.

CPS/USAR CHAPTER 02 2.5-116 REV. 11, JANUARY 2005 32. S. E. Harris, Jr. and M. C. Parker, "Stratigraphy of the Osage Series in Southeastern Iowa," Report of Investigations 1, 52 pp., Iowa Geological Survey, Iowa City, Iowa, 1964. 33. R. L. Brownfield, "Structural History of the Centralia Area," Report of Investigations No. 172, 31 pp., Illinois State Geological Survey, Urbana, Illinois, 1954. 34. A. V. Heyl, "The 38th Parallel Lineament and its Relationship to Ore Deposits,"

Economic Geology, Vol. 67, pp. 879-894, 1972. 35. Sargent & Lundy, Supplemental Discussion Concerning the Limit of the Northern Extent of Large Intensity Earthquakes Similar to the New Madrid Events, May 23, 1975. 36. J. A. Harrison, "Subsurface Geology and Coal Resources of the Pennsylvanian System in White County, Illinois," Report of Investigations No. 153, 34 pp., Illinois State Geological Survey, Urbana, Illinois, 1951. 37. J. M. Weller, R. M. Grogan, and F. E. Tippie, "Geology of the Fluorspar Deposits of Illinois," Bulletin 76, 147 pp., Illinois State Geological Survey, Urbana, Illinois, 1952. 38. J. W. Baxter and G. A. Desborough, "Areal Geology of the Illinois Fluorspar District, Part 2 - Karbers Ridge and Rosiclare Quadrangles," Circular 385, 40 pp., Illinois, State Geological Survey, Urbana, Illinois 1965. 39. W. N. Melhorn and N. M. Smith, "The Mt. Carmel Fault and Related Structural Features in South-Central Indiana," Report of Progress No. 16, 29 pp., Indiana Geological Survey, Bloomington, Indiana, 1959. 40. T. A. Dawson, "Map of Indiana Showing Structure on Top of Trenton Limestone," Miscellaneous Map 17, Indiana Geological Survey, Bloomington, Indiana, 1971. 41. L. E. Becker, Head, Petroleum Section, Indiana Geological Survey, written communication to A. K. Yonk, Sargent & Lundy Senior Geologist, May 22, 1975.

Attachment D2.5, CPS-FSAR. 42. H. H. Gray, Head Stratigrapher, Indiana Geological Survey, written communication to G. E. Heim, Sargent & Lundy Senior Geologist, October 15, 1974. Attachment D2.5, CPS-

FSAR. 43. M. E. Ostrom, Director and State Geologist, Wisconsin Geological and Natural History Survey, written communication to G. E. Heim, Sargent & Lundy Geotechnical Division

Head, April 8, 1975. Attachment D2.5, CPS-FSAR. 44. M. E. Ostrom, "Geology Field Trip - Southwestern Dane County," 10 pp., University of Wisconsin Geological and Natural History Survey, Madison, Wisconsin, 1971. 45. T. C. Buschbach and R. Ryan, "Ordovician Explosion Structure at Glasford, Illinois,"

Bulletin of the American Association of Petroleum Geologists, Vol. 47, No. 12, pp. 2015-2022, American Association of Petroleum Geologists, Tulsa, Oklahoma, 1963.

CPS/USAR CHAPTER 02 2.5-117 REV. 11, JANUARY 2005 46. G. H. Emrich and R. E. Bergstrom, " Des Plaines Disturbance, Northeastern, Illinois,"

Geological Society of America Bulletin, Vol. 73, pp. 959-968, Geological Society of America, Boulder, Colorado, 1962. 47. D. A. Green, "Trenton Structure in Ohio, Indiana, and Northern Illinois," Bulletin of the American Association of Petroleum Geologists, Vol. 41, pp. 627-642, American Association of Petroleum Geologists, Tulsa, Oklahoma, 1957. 48. T. C. Buschbach, Coordinator, Nuclear Facilities Siting Studies, Illinois State Geological Survey, written communication to Dr. T. Kemmis, Sargent & Lundy Geologist, April 25, 1977. Attachment D2.5, CPS-FSAR. 49. L. D. McGinnis, "Crustal Tectonics and Precambrian Basement in Northeastern Illinois," report of Investigations 219, 29 pp., Illinois, State Geological Survey, Urbana, Illinois

1966. 50. T. C. Buschbach and G. E. Heim, "Preliminary Geologic Investigations of Rock Tunnel Sites for Flood and Pollution Control in the Greater Chicago Area," Environmental Geology Note 52, 35 pp., Illinois State Geological Survey, Urbana, Illinois, 1972. 51. F. T. Thwaites, "Map of the Buried Pre-Cambrian of Wisconsin," Wisconsin Geological and Natural History Survey, Madison, Wisconsin, 1957. 52. L. D. McGinnis, et al., "The Gravity Field and Tectonics of Illinois," Circular 494, 24 pp., Illinois State Geological Survey, Urbana, Illinois, 1976. 53. D. C. Bond, et al., "Possible Future Petroleum Potential of Region 9 - Illinois Basin, Cincinnati Arch, and Northern Mississippi Embayment," AAPG Memoir 15, Future Petroleum Provinces of the United States - Their Geology and Potential, (Ed. by I. H. Cram, et al.), Vol. 2, p. 1165-1218, American Association of Petroleum Geologists, Tulsa, Oklahoma, 1971. 54. P. C. Heigold, "An Aeromagnetic Survey of Southwestern Illinois," Circular 495, 28 pp., Illinois State Geological Survey, Urbana, Illinois, 1976. 55. R. W. Patenaude, "Results of Regional Aeromagnetic Surveys of Eastern Upper Michigan, Central Lower Michigan, and Southeastern Illinois," Department of Geology Research Report 64-2, 51 pp., University of Wisconsin Geophysical and Polar Research Center, Madison, Wisconsin, April 1964. 56. L. D. McGinnis and P. C. Heigold, "Regional Maps of Vertical Magnetic Intensity in Illinois," Circular 324, 12 pp., Illinois State Geological Survey, Urbana, Illinois, 1961. 57. K. E. Clegg, "Subsurface Geology and Coal Resources of the Pennsylvanian System in DeWitt, McLean, and Piatt Counties, Illinois," Circular 473, Illinois State Geological Survey, Urbana, Illinois, 25 pp., 1972. 58. P. C. Heigold, L. D. McGinnis, and R. H. Howard, "Geologic Significance of the Gravity Field in the Dewitt, McLean County Area, Illinois," Circular 369, 16 pp., Illinois State Geological Survey, Urbana, Illinois, 1964.

CPS/USAR CHAPTER 02 2.5-118 REV. 11, JANUARY 2005 59. J. Van Den Berg and T. F. Lawry, "Petroleum Industry in Illinois, 1975" Illinois Petroleum 110, 126 pp., Illinois State Geological Survey, Urbana, Illinois, 1976. 60. T. C. Buschbach and D. C. Bond, "Underground Storage of Natural Gas in Illinois - 1973," Illinois Petroleum 101, 71 pp., Illinois State Geological Survey, Urbana, Illinois, 1974. 61. J. Docekal, "Earthquakes of the Stable Interior, with Emphasis on the Midcontinent," Ph.D. Dissertation, University of Nebraska, Lincoln, Vols. 1 and 2, 1970. 62. J.B. Hadley and J.F. Devine, "Seismotectonic Map of the Eastern United States," U.S. Geol. Survey. Misc. Field Studies, Map MF-620, 3 sheets, 1974. 63. A.J. Eardley, "Structural Geology of North America, Second Edition" Harper and Row, New York, 1962. 64. P.B. King, "The Tectonics of Middle North America," Hafner Publishing Company, New York, 1951. 65. M.L. Fuller, "The New Madrid Earthquake," Bulletin 494, U.S. Geol. Surv., 1912. 66. O.W. Nuttli, "The Mississippi Valley Earthquakes of 1811 and 1812, Intensities, Ground Motion, and Magnitudes," Seismol. Soc. Amer. Bull. 63, 1:227-248, 1973. 67. R.G. Stearns and C.W. Wilson "Relationship of Earthquakes and Geology in West Tennessee and Adjacent Areas," Tennessee Valley Authority, 1972. 68. L.D. McGinnis and C.P. Ervin, "Earthquakes and Block Tectonics in the Illinois Basin," Geology, 2:517-519, 1974. 69. Public Service Indiana, Marble Hill Preliminary Safety Analysis Report, Docket Nos. 50-546 and 50-547, 1975. 70. Stone & Webster, "Faulting in the Anna Ohio Region" Amendment 12, Appendix 2I PSAR for WUP, Koshkonong Station, 1976. 71. Union Electric Company, Callaway Preliminary Safety Analysis Report, Docket Nos. 50-488 and 50-486, 1976. 72. J.L. Coffman and C.A. von Hake, "Earthquake History of the United States," National Oceanic and Atmos. Adm., Boulder, Colo., Pub. 41-1 (revised edition through 1970),

1973. 73. N.S. Shaler, "Earthquakes of Western United States," Atlantic Monthly, 24(445): 549-559, 1869. 74. Sargent & Lundy, "Supplemental Discussion Concerning the Limit of the Northern Extent of Large Intensity Earthquakes Similar to the New Madrid Events," May 23, 1975. 75. W. Stauder and et al, "Seismic Characteristics of Southeast Missouri as Indicated by a Regional Telemetered Microearthquake Array," Seismol. Soc. Amer. Bull. 66, 6:1953-1964, December 1976.

CPS/USAR CHAPTER 02 2.5-119 REV. 11, JANUARY 2005 76. F.L. Fox and C.T. Spiker, "Intensity Rating of the Attica (N.Y.) Earthquake of August 12, 1929. A Proposed Reclassification," Earthquake Notes, Vol. 48, 1-2:37-46, Jan. - June 1977. 77. Kansas Gas and Electric Company/Kansas City Power & Light Company, Wolf Creek Preliminary Safety Analysis Report, Docket No. 50-482, 1974. 78. M.D. Trifunac and A.G. Brady, "On the Correlation of Seismic Intensity Scales with the Peaks of Recorded Strong Ground Motion," Seismol. Soc. Amer. Bull. 65, 1:139-162,

1975. 79(a). H. A. Merz and C. A. Cornell, "Seismic Risk Analysis Based on a Quadratic Magnitude-Frequency Law," Bulletin of the Seismological Society of America, Vol. 63, No. 6, pp. 1999-2006, December 1973. 79(b). C. A. Cornell and H. A. Merz, "Seismic Risk Analysis of Boston," Journal of the Structural Division, ASCE, Vol. 101, No. ST10, pp. 2027-2043, October 1975. 80. W. D. McClain and 0. H. Myers, "Seismic History and Seismicity of the Southeastern Region of the United States," Report No. ORNL-4582, Oak Ridge National Laboratory, Oak Ridge, Tennessee, June 1970. 81. I. N. Gupta, "Attenuation of Intensities Based on Isoseismals of Earthquakes in Central United States" Earthquake Notes, Vol. 47, No. 3, pp. 13-20, July-September 1976. 82. J. R. Murphy and L. J. O'Brien, "Analysis of a Worldwide Strong Motion Data Sample to Develop an Improved Correlation Between Peak Acceleration, Seismic Intensity, and Other Physical Parameters," NUREG-0402, Computer Sciences Corporation (Report prepared for USNRC), January 1978. 83. L. Bjerrum, "Problems of Soil Mechanics and Construction on Soft Clays," Proceedings of 8th International Conference on Soil Mechanics and Foundation Engineering, Moscow, U.S.S.R., 1973. 84. "Standard Methods for the Examination of Water and Wastewater", 13th Edition, American Public Health Association, et al., Washington, D.C., 1971. 85. H. B. Seed and J. M. Idriss, "Simplified Procedure for Evaluating Soil Liquefaction Potential," Journal of Soil Mechanics and Foundations Division, ASCE, 97(SM9): pp.

1249-1273, 1971. 86. H. B. Seed and W. H. Peacock, "Test Procedures for Measuring Soil Liquefaction Characteristics," Journal of Soil Mechanics and Foundations Division, ASCE, 97 (SM8):

pp 1099-1119, August 1971. 87. A. Casagrande, "The Determination of Preconsolidation Load and its Practical Significance," in Proceedings of First International Conference on Soil Mechanics and Foundation Engineering, Cambridge, Massachusetts, 1936. 88. "Design of Small Dams", U.S. Department of the Interior, Bureau of Reclamation, U.S. Government Printing Office, Washington, D.C., 1965.

CPS/USAR CHAPTER 02 2.5-120 REV. 11, JANUARY 2005 89. "Engineering and Design of Earth- and Rock-Fill Dams: General Design and Construction Considerations," Engineer Manual EM 110-2-2300, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, 1971. 90. J. S. Templeton and H. B. Willman, "Guidebook for the Sixteenth Annual Field Conference of the Tri-State Geological Society," Guidebook Series 2, Illinois State Geological Survey, Urbana, Illinois, 47 pp., 1952. 91. G. V. Cohee and C. W. Carter, "Structural Trends in the Illinois Basin," Circular 59, Illinois State Geological Survey, Urbana, Illinois, 4 pp., 1940. 92. K. E. Clegg, "Subsurface Geology and Coal Resources of the Pennsylvanian System in Clark and Edgar Counties, Illinois," Circular 380, Illinois State Geological Survey, Urbana, Illinois, 28 pp., 1965. 93. H. M. Bristol and R. Prescott, "Geology and Oil Production in the Tuscola Area, Illinois," Circular 424, Illinois State Geological Survey, Urbana, Illinois, 34 pp., 1968. 94. H. B. Seed and W. H. Peacock, "Applicability of Laboratory Test Procedures for Measuring Soil Liquefaction Characteristics Under Cyclic Loading," Report No. EERC 70-8, Earthquake Engineering Research Center, University of California, Berkeley, California, 1970. 95. H. B. Seed and R. V. Whitman, "Design of Earthquake Retaining Structures for Dynamic Loads, ASCE Specialty Conference on Lateral Stresses in the Ground and Design of Earth Retaining Structures," Cornell University, Ithaca, N.Y., 1970. 96. J. P. Gould, "Lateral Pressures on Rigid Permanent Structures," ASCE Specialty Conference on Lateral Earth Pressures, Cornell University, 1970. 97. G. F. Sowers et al., "The Residual Lateral Pressures Produced by Compacting Soils," Fourth Interior Conference, SM&FE, Volume II, London, 1957. 98. K. L. Lee and H. B. Seed, "Cyclic Stress Condition Causing Liquefaction of Sand," Journal of Soil Mechanics and Foundations Division, ASCE, (SMI), January 1967. 99. H. B. Seed, "Landslides during Earthquakes due to Soil Liquefaction," Journal of the Soil Mechanics and Foundations Division, 94(SM5), September 1968. 100. I. M. Idriss and Y. Lacroix, "Preliminary Safety Analysis Report," Appendix 2E, Shearon Harris Nuclear Power Plant, Wake County, North Carolina, Carolina Power and Light Company, 1973. 101. H. B. Seed and C. K. Chan, "Clay Strength under Earthquake Loading Conditions," Journal of Soil Mechanics and Foundations Division, ASCE, 92(SM2): 4723-4733, 1966. 102. L. D. McGinnis, P. C. Heigold, C. P. Ervin and M. Heidari, "The Gravity Field and Tectonics of Illinois," Illinois State Geological Survey Circular 494, 28 pp., 1976. 103. P. C. Heigold, "An Aeromagnetic Survey of Southwestern Illinois," Illinois State Geological Survey Circular 495, 28 pp., 1976.

CPS/USAR CHAPTER 02 2.5-121 REV. 11, JANUARY 2005 104. U. S. Nuclear Regulatory Commission, Office of Nuclear Reactor Regulation, "Safety Evaluation Report Related to Construction of Marble Hill Plant," Docket Nos. 50-546 and 50-547, 1977. 105. H. B. Seed, K. L. Lee, and I. M. Idriss, "Analysis of Sheffield Dam Failure," Journal of Soil Mechanics and Foundations Division, ASCE, 95(SM6): 1453-1490, 1969. 106. F. H. Kulhawy, J. M. Duncan, and H. B. Seed, "Finite Element Analyses of Stresses and Movements in Embankments during Construction," College of Engineering, Office of Research Services, University of California, Berkeley, California, Report Number TE 4, 1969. 107. J. M. Roesset, "Fundamental of Soil Amplification, Seismic Design for Nuclear Power Plants," Edited by R. J. Hansen, The M.I.T. Press, Cambridge, Massachusetts, 1970. 108. H. B. Seed et al., "Analysis of the Slides in the San Fernando Dams During the Earthquake of February 9, 1971," Report No. EERC 73-2, Earthquake Engineering Research Center, University of California, Berkeley, 1973. 109. W. L. Shannon and D. E. Hiltz, "Submarine Landslide at Seward, The Great Alaska Earthquake of 1964." 110. R. L. Shiffman, R. V. Whitman, and J. C. Jordan, "Settlement Problem Oriented Computer Language," Journal of Soil Mechanics and Foundations Division, ASCE,

March 1970. 111. McDonnell Douglas Automation Company, "SEPOL (User's Manual)," 1972. 112. K. Terzaghi and R. B. Peck, "Soil Mechanics in Engineering Practice," John Wiley & Sons, Inc., New York, 1968. 113. H. B. Seed and W. H. Peacock, "Applicability of Laboratory Test Procedures for Measuring Soil Liquefaction Characteristics under Cyclic Loading," Report No. EERC 70-8, Earthquake Engineering Research Center, University of California, Berkeley, 1970. 114. H. B. Seed, "Landslides During Earthquakes Due to Liquefaction," Journal of the Soil Mechanics and Foundations Division, ASCE, Volume 94, No. SM5, Proceeding Paper 6110, pp. 1053-1122, September 1968. 115. E. D'Appolonia, "Dynamic Loadings," Journal of Soil Mechanics and Foundations Division, ASCE, No. SM1, Volume 96, p. 61, January 1970. 116. Site Specific Response Spectra for Clinton Power Station - Unit 1 of Illinois Power Company, Revision 0, February 1982, by Weston Geophysical Corporation.

CPS/USAR CHAPTER 02 2.5-122 REV. 11, JANUARY 2005 TABLE 2.5-1 SUMMARY OF FOLDS NAME AND STATE MEANS OF IDENTIFICATION* MAJOR MOVEMENT** Illinois Ashton Arch B,S Late Paleozoic (Reference 90)Cap au Gres Faulted Flexure S, B Post-Middle Mississippian, Pre-Pennsylvanian (Reference 26) Clay City Anticline B Pre-Pennsylvanian, Pennsylvanian, and/or Post-

Pennsylvanian (Reference 25)

Downs Anticline B Mississippian through Pennsylvanian (Reference 57)Dupo-Waterloo Anticline B,S Late Missisippian, Pre-Pennsylvanian (Reference 22)

DuQuoin Monocline B Pennsylvanian or earlier (Reference 22) Illinois Basin S, B, G Early to Late Paleozoic (Reference 2) Kankakee Arch S, B, G Ordovician to Pennsylvanian (Reference 15) La Salle Anticlinal Belt S, B, G Post-Mississippian to Permian (Reference 22) Lincoln Anticline S,B Late-Mississippian to Early Pennsylvanian (Reference 23)

Marshall Syncline B Late or Post-Pennsylvanian (Reference 92)

Matoon Anticline B Late Paleozoic (Reference 92)Mississippi River Arch S, B, G Late Paleozoic (Reference 22)

Moorman Syncline S, B Post-Pennsylvanian (Reference 20)

Murdock Syncline B Late or Post-Pennsylvanian (Reference 92) Pittsfield-Hadley Anticline B Post Pennsylvanian (Reference 22) Salem and Loudon Anticlines B Pennsylvanian and Post-Pennsylvanian (Reference 22)Sangamon Arch B, G Devonian to Early Mississippian (Reference 22)

CPS/USAR TABLE 2.5-1 (Cont'd) CHAPTER 02 2.5-123 REV. 11, JANUARY 2005 NAME AND STATE MEANS OF IDENTIFICATION*

MAJOR MOVEMENT** Illinois (Cont'd)

Structures associated with the Plum River Fault Zone S, B, G Pennsylvanian, Post-Pennsylvanian, Pre-Middle Illinoian (Pleistocene)

(Reference 28) Tuscola Anticline B Pennsylvanian, and Post-Pennsylvanian (Reference 93)

Iowa Bentonsport B Mississippian (Reference 32)

Burlington B Mississippian (Reference 32)

Oquawka B Mississippian (Reference 32)

Skunk River B Mississippian (Reference 32)

Sperry B Mississippian (Reference 32)

Missouri Auxvasse Creek Anticline S Post-Pennsylvanian (Reference 17) Browns Station Anticline S Late Mississippian or Pennsylvanian (Reference

Clinton PSAR)

College Mound-Bucklin

Anticline S Later Part or Post-Pennsylvanian (Reference 17)Crystal City Anticline S Post-Mississipian (Reference Clinton PSAR)

Cuivre Anticline S Post-Mississippian (Reference Clinton PSAR)

Davis Creek Anticline B Post-Mississippian (Reference Clinton PSAR)

Eureka-House Springs

Anticline S, B Post-Mississippian (Reference

17) Farmington Anticline S, B No older than Devonian (Reference 17)

Kruegers Ford Anticline S, B Post-Ordovician (Reference Clinton PSAR)

CPS/USAR TABLE 2.5-1 (Cont'd) CHAPTER 02 2.5-124 REV. 11, JANUARY 2005 NAME AND STATE MEANS OF IDENTIFICATION*

MAJOR MOVEMENT**

Missouri (Cont'd)

Mexico Anticline S, B Late or Post-Pennsylvanian (Reference Clinton PSAR)

Mineola Structure S Pennsylvanian, Post-Pennsylvanian (Reference Clinton PSAR) Ozark Uplift S, B, G Paleozoic, Mesozoic, Tertiary (Reference 17) Pershing-Bay-Gerald Anticline S Post-Mississippian, Early Pennsylvanian (Reference

Clinton PSAR) Plattin Creek Anticline S Post-Mississippian (Reference Clinton PSAR)

Troy Brussels Syncline S, B Late Mississippian or Early Pennsylvanian (Reference 26)

Wisconsin Meekers Grove Anticline B, S Late Paleozoic (Reference 31)

Mineral Point Anticline B, S Late Paleozoic (Reference 31)Wisconsin Arch S, B, G Early to Late Paleozoic (Reference 15)

__________________________ *S = surface mapping, B = borehole, G = geophysical

• • Due to the absence of sediments representing the interval from Pennsylvanian to Cretaceous or Pleistocene time, the age of final movement on these structures cannot be precisely dated. However, based on stratigraphic relationships and geologic history outside of the regional area, final movement on the structures is considered to have occurred prior to Pleistocene time, and possibly before the end of the Paleozoic.

CPS/USAR CHAPTER 02 2.5-125 REV. 11, JANUARY 2005 TABLE 2.5-2 SUMMARY OF FAULTS FAULT NAME MEANS OF IDENTIFICATION*FAULT DISPLACEMENT LAST MOVEMENT Cap au Gres Faulted Flexure S, B Maximum structural relief of 1200 feet (Reference 26) Late Plicene-Pre-Pleistocene

(Reference 22) Centralia Fault S (in mines), B Downthrown up to 200 feet on west side (Reference 22)

Post-Pennsylvanian, Pre-Pleistocene (Reference 22)

Chicago Area

Basement Fault**

(Inferred)

G, B Downthrown 900 feet on southwest side (Reference 49)

Pre-Middle Ordovician (Reference

49) Chicago Area Minor Faults (Inferred)

G North or south side of faults downthrown, 55 feet max.

displacement (Reference 50) Post-Ordovician or Post-Silurian, Pre-Pleistocene (Reference 50)

Chicago Area Minor Faults S Few inches to Few Feet (Reference 50) Post-Silurian, Pre-Pleistocene

(Reference 50) Fortville Fault B Downthrown 60 feet on southeast side (Reference 41)

Post-Devonian, Pre-Pleistocene

(Reference 42)

(Janesville

Fault)***,+ (Iniferred)

B North side downthrown (Reference 51)

Phanerozioc But, Pre-Pleistocene

possible Pre-Cretaceous

(Reference 43, 51)

(Madison Fault)***,+

(Inferred)

B North side downthrown (Reference 51)

Phanerozoic but Pre-Pleistocene, possibly Pre-Cretaceous (References 43 and 51) Mt. Carmel Fault S, B Downthrown in excess of 200 feet on west side (Reference 39) Early Pennsylvanian Reference

39)

CPS/USAR TABLE 2.5-2 (Cont'd) CHAPTER 02 2.5-126 REV. 11, JANUARY 2005 FAULT NAME MEANS OF IDENTIFICATION*FAULT DISPLACEMENT LAST MOVEMENT Northeast Trending Faults South of the

Rough Creek Fault

Zone S, B Graben structures present, northwest or southeast walls of faults downthrown.

Displacements are variable, may be up to 2000 feet. (Reference 38)

Tertiary, possibly Pre-Cretaceous

(Reference 14)

Oglesby Fault+

(Inferred) B Downthrown 1200 feet on west side (Reference 47)

Pre-Cretaceous (Reference

Clinton PSAR) Plum River Fault

Zone S, B, G Downthrown up to 400 feet on north side (Reference 28) Post-Silurian, Pre-Middle Illinoian

(Reference 28)

Rough Creek Fault

Zone S, B, G North side downthrown; max.

reported displacement of 3400 feet along the fault zone.

(Reference 22) Post-Pennsylvanian Pre-Late Cretaceous (Reference 22) Royal Center Fault B Downthrown 100 feet on southeast side (Reference 41)

Post-Middle Devonian, Pre-Pleistocene (Reference 42) Ste. Genevieve Fault Zone S, B, G Net displacement along the fault zone is down to the north and east; max. displacement greater than 1000 feet, possibly as much as 2000 feet (Reference 22)

Post-Pennsylvanian, Pre-Pleistocene (Reference 20,

Reference 2) Sandwich Fault Zone S, B, G Downthrown up to 800 feet on northeast side (Reference 20) Post-Pennsylvanian Pre-Pleistocene (Reference 30) Tuscola Fault+

(Inferred) B Downthrown 2000 feet on west side (Reference 47)

Pre-Cretaceous (Reference

Clinton PSAR)

Wabash Valley Fault

Zone B, G, S Graben structures present; northwest or southeast sides of faults downthrown; max. displacement of 400 feet

(Reference 22)

Post-Pennsylvanian, Pre-Pleistocene (Reference 22)

CPS/USAR TABLE 2.5-2 (Cont'd) CHAPTER 02 2.5-127 REV. 11, JANUARY 2005 FAULT NAME MEANS OF IDENTIFICATION*FAULT DISPLACEMENT LAST MOVEMENT (Waukesha Fault)***,+,f Inferred) S, B Downthrown 1500+ feet on southeast side (Reference 51)

Phanerozoic but Pre-Pleistocene, possibly Pre-Cretaceous (References 43 and 51)

Wisconsin Minor Faults: Dane County S, B Downthrown up to 300 feet on the northwest side (Reference 44) Post-Ordovician, Pre-Pleistocene, possibly Pre-Cretaceous (References 43 and 44)

Wisconsin Minor Faults: Waukesha f S Downthrown 27 feet on southeast side (Reference 43) Post-Silurian Pre-Pleistocene, possibly Pre-Cretaceous (Reference 43) and Geological Map of the U.S., U.S. Geological Survey, (1974)

___________________________

Notes:

• S = surface, B = borehole, G = geophysical

• • No confirmation has been made for this postulated fault.

• • • Names in parentheses were assigned by Dames & Moore.

+ Recent authorities doubt the existence of these faults, see References 43 and 48 which also appear in Attachment D2.5.

f For the present interpretation of these faults see Reference 43, which also appears in Attachment D2.5.

CPS/USAR CHAPTER 02 2.5-128 REV. 11, JANUARY 2005 TABLE 2.5-3 CHEMICAL ANALYSIS ON THE NO. 7 AND NO. 8 COAL FROM BORING P-38 BY THE ILLINOIS STATE GEOLOGICAL SURVEY SAMPLE CONDITION* MOISTURE

(%) VOLATILE MATTER (%)

FIXED CARBON (%)

ASH (%) TOTAL SULFUR (%)

BTU/LBNo. 8 coal 1 12.5 32.2 35.9 19.3 4.2 9,600 2 -- 36.8 41.1 22.1 4.8 10,900 3 -- 47.2 52.8 -- 6.1 14,000 4 16.3 37.7 46.0 -- -- 12,200 5 -- 45.1 54.9 -- -- 14,600

No. 7 coal 1 9.6 35.4 35.8 19.3 8.2 9,800 2 -- 39.1 39.6 21.3 9.1 10,800 3 -- 49.7 50.3 -- 11.5 13,800 4 12.8 40.9 46.3 -- -- 12,600 5 -- 46.9 53.1 -- -- 14,400

_________________________

• Type of analyses noted as follows: 1. as received at laboratory,

2. moisture free,

3. moisture and ash free,

4. moist mineral matter free,

5. dry mineral matter free.

CPS/USAR CHAPTER 02 2.5-129 REV. 11, JANUARY 2005 TABLE 2.5-4 EARTHQUAKE EPICENTERS, 37° to 45° NORTH LATITUDE, 84° to 93° WEST LONGITUDE DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY

(MM) FELT AREA (mi²) REFERENCES 1. 1795 Jan 8 Franklin County, IL 37.9 89.0 IV-V 1 2. 1804 Aug 24 Fort

Dearborn,

IL 42.0 87.8 VI 30,000 1, 2, 3, 5 3. 1818 April 11 St. Louis, MO 38.6 90.2 III-IV 7,500 1 4. 1819 Sept 2 Perry County, MO 37.7 89.7 IV 15,000 2 5. 1819 Sept 16 Randolph County, IL 38.1 89.8 IV 9,600 1 6. 1819 Sept 17 Randolph County, IL 38.1 89.8 III-IV 1 7. 1820 Nov 9 Cape Girardeau, MO 37.3 89.5 IV 5,000 2, 3 8. 1827 July 5 St. Louis, MO 38.6 90.2 IV-V 1 9. 1827 July 5 Grant County, KY 38.7 84.6 IV 15,000 1, 2 10. 1827 July 5 New Albany, IN 38.3 85.8 165,000 1, 2 11. 1827 July 6 Cincinnati, OH 39.1 84.5 IV 1 12. 1827 Aug 6 New Albany, IN 38.3 85.8 VI 1, 2, 3, 5 13. 1827 Aug 7 New Albany, IN 38.3 85.8 VI 1, 2, 3, 5 14. 1827 Aug 14 St. Louis, MO 38.6 90.2 III 1 15. 1838 June 9 St. Louis, MO 38.6 90.2 VI 300 1 16. 1843 Feb 16 St. Louis, MO 38.6 90.2 IV-V 100,000 1 17. 1845 Putnam County, OH 41.1 84.2 II 1 18. 1850 April 4 Louisville, KY 38.3 85.8 IV 1, 2, 4 19. 1854 Feb 28 Lexington, KY 38.1 84.5 VI 15 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-130 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 20. 1854 Feb 28 Garrard County, KY 37.6 84.5 IV 8,000 2 21. 1855 May 2 Cairo, IL 37.0 89.2 IV 2 22. 1855 May 3 Cairo, IL 37.0 89.2 III 2 23. 1857 Oct 8 St. Louis, MO 38.6 90.3 VI-VII 7,500 1, 3 24. 1860 Aug 7 Henderson, KY 37.8 87.6 V 30,000 2 25. 1869 Feb 20 Lexington, KY 38.1 84.5 III-IV 1 26. 1871 July 24 Cairo, IL 37.0 89.2 III 2 27. 1871 July 25 St. Clair County, IL 38.5 90.0 III 1,000 1 28. 1872 Feb 8 Cairo, IL 37.0 89.2 III 1 29. 1872 March 26 Paducah, KY 37.1 88.6 III 2 30. 1873 April 22 Dayton, OH 39.8 84.2 III-IV 1 31. 1874 July 9 Cairo, IL 37.0 89.2 III-IV 2 32. 1875 June 18 Champaign County, OH 40.2 84.0 VII 40,000 1,2,3,6,8,14 33. 1876 Jan 27 Adrian, MI 41.9 84.0 NOT RECORDED 1 34. 1876 June Anna, OH 40.4 84.2 I-III 1,8,14,19 35. 1876 Sept 24 Wabash County, IL 38.5 87.9 VI 1 36. 1876 Sept 25 Knox County, IN 38.5 87.7 VI 60,000 1,2,3,6,7 37. 1876 Sept 26 Wabash County, IL 38.5 87.9 III 1 38. 1877 May 26 New Harmony, IN 38.1 87.9 III-IV 1 39. 1877 June 3 Stanford, KY 37.5 84.7 III 2 40. 1877 July 15 Carbondale, IL 37.7 89.2 III-IV 9,500 2 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-131 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 41. 1878 Jan 8 Cairo, IL 37.0 89.2 III-IV 3,000 2 42. 1878 Nov 19 Cairo, IL 37.0 89.2 III 3,000 2 43. 1881 April 20 Goshen, IN 41.6 85.8 IV 1, 2 44. 1881 May 27 LaSalle, IL 41.3 89.1 VI 1, 2 45. 1881 Aug 29 Hillsboro, OH 39.2 83.6 III 1, 2 46. 1882 Feb 9 Anna, OH 40.4 84.2 V 100 1, 2, 3, 8, 14 47. 1882 July 20 Randolph County, IL 38.0 90.0 V 30,000 1, 2 48. 1882 Sept 27 Macoupin County, IL 39.0 90.0 VI 25,000 1, 2, 3 49. 1882 Oct 14 Macoupin County, IL 39.0 90.0 V 8,000 1, 2 50. 1882 Oct 15 Macoupin County, IL 39.0 90.0 V 8,000 1, 2, 3 51. 1882 Oct 22 Greenville, IL 38.9 89.4 III 1 52. 1882 Nov 15 St. Louis, MO 38.6 90.2 III 1 53. 1883 Jan 10 Union County, IL 37.5 89.3 III 2 54. 1883 Jan 11 Cairo, IL 37.0 89.2 VI 80,000 3 55. 1883 Feb 4 Kalamazoo County, MI 42.3 85.6 VI 150,000 1, 2, 3 56. 1883 April 12 Cairo, IL 37.0 89.2 VI-VII 2 57. 1883 July 6 Cairo, IL 37.0 89.2 III 2 58. 1883 July 14 Wickliffe, KY 37.0 89.1 IV-V 10,000 2 59. 1883 Nov 14 St. Louis, MO 38.6 90.2 IV 1,200 1 60. 1883 Dec 28 Bloomington, IL 40.5 87.0 III 16 61. 1884 Feb 15 Washington County, MO 37.8 90.8 III 2 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-132 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 62. 1884 March 31 Preble County, OH 39.6 84.8 II 1 63. 1884 Sept 19 Allen County, OH 40.7 84.1 V-VI 125,000 1, 2, 8, 14 64. 1884 Dec 23 Anna, OH 40.4 84.2 III 1, 5, 14 65. 1885 Dec 26 Bloomington, IL 40.5 89.0 III 1 66. 1886 March 1 Butlerville, IN 39.0 85.5 III-IV 1,2 67. 1886 March 18 Cairo, IL 37.0 89.2 III-IV 3,000 2 68. 1886 Aug 13 Indianapolis, IN 39.8 86.2 III-IV 1 69. 1887 Feb 6 Vincennes, IN 38.7 87.4 VI 75,000 1, 2, 3, 6, 7 70. 1887 Aug 2 Cairo, IL 37.0 89.2 V 65,000 2 71. 1889 Sept Anna, OH 40.4 84.2 III 1, 8, 14 72. 1891 July 26 Evansville, IN 38.0 87.6 VI 1, 2, 3, 6 73. 1891 Sept 26 Cairo, IL 37.0 89.2 V 2 74. 1892 Anna, OH 40.4 84.2 IV-VI 1,8,14,19 75. 1895 Oct 31 Near Charleston, MO 37.0 89.4 VIII 1,000,000 2, 3 76. 1895 Nov 2 Near Charleston, MO 37.0 89.4 III-IV 1 77. 1895 Nov 17 Near Charleston, MO 37.0 89.4 III-IV 1 78. 1896 March 15 Sidney, OH 40.3 84.2 IV 1, 8, 14 79. 1897 Oct 31 Niles, MI 41.8 86.3 NOT RECORDED 1 80. 1898 June 6 Richmond, KY 37.8 84.3 III 2 81. 1899 Feb 8 Chicago, IL 41.9 87.6 IV-V 1 82. 1899 Feb 9 Chicago, IL 41.9 87.6 IV-V 1 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-133 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 83. 1899 April 29 Dubois County, IN 38.5 87.0 VI-VII 40,000 1, 2, 6 84. 1899 Oct 10 St. Joseph, MI 42.1 86.5 IV 1 85. 1899 Oct 12 Kenosha, WI 42.6 87.8 NOT RECORDED 1 86. 1902 Jan 24 Maplewood, MO 38.6 90.3 VI 40,000 1, 3 87. 1902 March 10 Hagerstown, IN 39.9 85.2 III-IV 1 88. 1903 Jan 1 Hagerstown, IN 39.9 85.2 II-III 1 89. 1903 Feb 8 St. Louis, MO 38.6 90.3 VI 40,000 1, 3 90. 1903 March 17 Hillsboro, IL 39.2 89.5 III-IV 1 91. 1903 Sept 20 Morgantown, IN 39.4 86.3 IV 1 92. 1903 Sept 21 Olney, IL 38.7 88.1 IV 1 93. 1903 Nov 3 Murphysboro, IL 37.8 89.3 III-IV 2 94. 1903 Nov 4 St. Louis, MO 38.6 90.3 VII 70,000 Callaway PSAR 95. 1903 Nov 20 Morgantown, IN 39.4 86.3 1 96. 1903 Dec 11 Effingham, IL 39.1 88.5 II 1 97. 1903 Dec 31 Fairmont, IL 41.6 88.1 1 98. 1905 March 13 Menominee, MI 45.0 87.7 V 1, 3 99. 1905 April 13 Keokuk, IA 40.4 91.6 V 5,000 1, 2, 3 100. 1905 Aug 22 Quincy, IL 39.9 91.4 II-III 1 101. 1906 Feb 23 Anabel, MO 39.7 92.4 III 1 102. 1906 March 6 Hannibal, MO 39.7 91.4 IV 1 103. 1906 April 22 Milwaukee, WI 43.0 87.9 1 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-134 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 104. 1906 April 24 Milwaukee, WI 43.0 87.9 1 105. 1906 May 8 Shelby County, IN 39.5 85.8 III-IV 600 1 106. 1906 May 9 Columbus, IN 39.2 85.9 IV 1, 2, 3 107. 1906 May 11 Petersburg, IN 38.5 87.3 V 1,200 1, 2, 3 108. 1906 May 19 Grand Rapids, MI 43.0 85.7 1 109. 1906 May 21 Flora, IL 38.7 88.5 V 560 1, 2, 3, 6 110. 1906 Aug 13 Greencastle, IN 39.6 86.9 IV 1 111. 1906 Sept 7 Owensville, IN 38.3 87.7 IV 500 1 112. 1906 Nov 23 Anabel, MO 39.7 92.4 III 1 113. 1907 Jan 10 Menominee, MI 45.1 87.6 1 114. 1907 Jan 29 Morgan County, IN 39.5 86.6 V 1, 2 115. 1907 Jan 30 Greenville, IL 38.9 89.4 V 1 116. 1907 July 4 Farmington, MO 37.8 90.4 IV-V 400 2, 3 117. 1907 Nov 20 Stephenson County, IL 42.3 89.8 IV 100 1, 2 118. 1907 Nov 28 Stephenson County, IL 42.3 89.8 IV 100 1, 2 119. 1907 Dec 10 St. Louis, MO 38.6 90.2 IV 1 120. 1908 Oct 27 Cairo, IL 37.0 89.2 V 5,000 2, 3 121. 1908 Dec 27 Ballard County, KY 37.0 89.0 IV 30,000 2 122. 1908 Dec 31 Ballard County, KY 37.0 89.0 III 1 123. 1909 May 26 South Beloit, IL 42.5 89.0 VII 170,000 1, 2, 3, 5 124. 1909 July 18 Mason County, IL 40.2 90.0 VII 35,000 1, 2, 3 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-135 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 125. 1909 Aug 16 Monroe County, IL 38.3 90.2 IV-V 18,000 1 126. 1909 Sept 22 Lawrence County, IN 38.7 86.5 V 4,000 1, 2, 3 127. 1909 Sept 27 Robinson, IL 39.0 87.7 VII 30,000 1, 2, 3, 6, 10 128. 1909 Sept 27 Vincennes, IN 38.7 87.5 V 4,000 1, 2, 3, 6, 10 129. 1909 Oct 22 Ironton, MO 37.6 90.6 IV 2 130. 1909 Oct 22 Sterling, IL 41.8 89.7 IV-V 1, 2 131. 1909 Oct 22 Near Scott, KY 38.9 84.5 IV or less 1 132. 1909 Oct 23 Scott County, MO 37.0 89.5 V 40,000 2, 3 133. 1909 Oct 23 Robinson, IL 39.0 87.7 V 14,000 1, 2, 5 134. 1911 Feb 28 St. Louis County, MO 38.7 90.3 IV 1 135. 1911 July 29 Chicago, IL 41.9 87.6 IV-V 1, 2 136. 1912 Jan 2 Kendall County, IL 41.5 88.5 VI 40,000 1, 3 137. 1912 Sept 25 Rockford, IL 42.3 89.1 III-IV 1, 2 138. 1913 Oct 16 Sterling, IL 41.8 89.7 III-IV 4,000 1, 2 139. 1913 Nov 11 Louisville, KY 38.3 85.8 IV 1 140. 1914 Oct 7 Madison, WI 43.1 89.4 IV 1 141. 1914 Anna, OH 40.4 84.2 III 1, 8, 14 142. 1915 Feb 5 Harrisburg, IL 37.7 88.5 IV 400 2 143. 1915 Feb 18 Mound City, IL 37.1 89.2 IV 350 2 144. 1915 April 15 Olney, IL 38.7 88.1 II-III 3,000 1 145. 1916 Jan 7 Worthington, IN 39.1 87.0 III 3,000 1 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-136 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 146. 1916 Feb 17 Near New Burnside, IL 37.6 88.8 III 2 147. 1916 May 31 Madison, WI 43.1 89.4 II 1 148. 1916 Aug 24 Cairo, IL 37.0 89.2 IV 4,000 2 149. 1916 Clarke County, IA 41.1 93.8 II-III 1 150. 1917 April 9 Jefferson County, MO 38.1 90.6 VI 200,000 1, 3 151. 1918 Feb 17 Cairo, IL 37.0 89.2 III 3,000 2 152. 1918 Feb 22 Shiawassee County, MI 42.9 84.2 IV 1 153. 1918 July 1 Hannibal, MO 39.7 91.4 IV 1 154. 1919 Feb 10 Henderson County, KY 37.8 87.5 III-IV 2,000 2 155. 1919 May 25 Knox County, IN 38.5 87.5 V 18,000 1, 2, 3, 6 156. 1920 April 30 Centralia, IL 38.5 89.1 IV 4,000 1 157. 1920 May 1 St. Louis County, MO 38.5 90.5 V 10,000 1, 3 158. 1921 Feb 27 Cairo, IL 37.0 89.2 III 3,000 2 159. 1921 March 14 Crawfordsville, IN 40.0 86.9 IV 25,000 1 160. 1921 March 31 Mount Vernon, IN 37.9 87.9 IV 2 161. 1921 Sept 8 Waterloo, IL 38.3 90.2 IV 4,000 1 162. 1921 Oct 1 Harrisburg, IL 37.7 88.5 IV 4,000 2 163. 1921 Oct 9 Waterloo, IL 38.3 90.2 III 3,000 1 164. 1922 Jan 10 Mount Vernon, IN 37.9 87.9 IV-V 9,500 2 165. 1922 March 22 Pope County, IL 37.3 88.6 V 25,000 3 166. 1922 March 22 Massac County, IL 37.3 88.9 V 60,000 2 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-137 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 167. 1922 March 22 Massac County, IL 37.3 88.9 V 2 168. 1922 March 23 Ballard County, KY 37.0 88.9 V 20,000 2 169. 1922 April 10 Monmouth, IL 40.9 90.7 II 1 170. 1922 July 7 Fond du Lac, WI 43.8 88.5 V 1, 2 171. 1922 Nov 26 Eldorado, IL 37.8 88.4 VI-VII 50,000 2 172. 1923 March 8 Greenville, IL 38.9 89.4 III-IV 4,000 1 173. 1923 May 6 Cairo, IL 37.0 89.2 III-IV 4,000 2 174. 1923 Nov 9 Tallula, IL 40.0 89.9 IV-V 600 1, 2, 3 175. 1923 Nov 28 Calhoun, KY 37.5 87.3 III 2 176. 1923 Nov 29 Mississippi County, MO 37.0 89.2 IV 2 177. 1924 April 2 Paducah, KY 37.1 88.6 IV 2 178. 1925 Jan 26 Waterloo, IA 42.5 92.3 II 200 1 179. 1925 March 3 Evanston, IL 42.0 87.7 II-III 1 180. 1925 April 4 Cincinnati, OH 39.1 84.5 IV or less 1, 8, 14 181. 1925 April 26 Vanderburgh County, IN 38.0 87.5 V-VI 100,000 1, 2, 3 182. 1925 July 13 Edwardsville, IL 38.8 90.0 V 1 183. 1925 Sept 2 Henderson County, KY 37.8 87.6 V-VI 75,000 2, 3 184. 1925 Sept 20 Henderson County, KY 37.8 87.6 IV 9,500 2 185. 1925 Oct Anna, OH 40.4 84.2 III 1, 8, 14 186. 1926 March 22 Harrisburg, IL 37.7 88.5 IV 4,000 2 187. 1926 Oct 3 Princeton, IN 38.4 87.6 III 1 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-138 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 188. 1927 Jan 31 Jackson, MO 37.4 89.7 IV 4,000 2 189. 1928 Jan 23 Near Mt. Carroll, IL 42.0 90.0 IV 400 1, 2 190. 1928 March 17 St. Louis, MO 38.6 90.2 I 1 191. 1928 April 15 Near Cape Girardeau, MO 37.4 89.7 III-IV 1 192. 1928 Oct 27 Shelby County, OH 40.4 84.1 III 100 1, 8, 14 193. 1929 Feb 14 Near Princeton, IN 38.3 87.6 III-IV 1,000 1 194. 1929 Feb 26 Arcadia, MO 37.5 90.6 III-IV 2 195. 1929 March 8 Shelby County, OH 40.4 84.2 V 5,000 1, 2, 3, 6, 8, 14 196. 1930 Feb 25 Near Cairo, IL 37.0 89.5 III 2 197. 1930 May 28 Near Hannibal, MO 39.7 91.3 III 1 198. 1930 June 26 Near Lima, OH 40.5 84.0 III-IV 1, 8, 14 199. 1930 June 27 Near Lima, OH 40.5 84.0 IV 1, 8, 14 200. 1930 Aug 8 Near Hannibal, MO 39.6 91.4 III-IV 1 201. 1930 Aug 29 Near Blandville, KY 37.0 89.2 IV 4,000 2 202. 1930 Sept 20 Anna, OH 40.4 84.2 VI 1, 2, 3, 8, 11, 14 203. 1930 Sept 29 Sidney, OH 40.3 84.2 III 1, 8, 14 204. 1930 Sept 30 Anna, OH 40.3 84.3 VII 1, 2, 3, 8, 9, 14 205. 1930 Oct Anna, OH 40.4 84.2 III-IV 1, 8, 14 206. 1930 Dec 23 Near St. Louis, MO 38.6 90.5 III-IV 1,000 1 207. 1931 Jan 5 Elliston, IN 39.0 86.9 V 500 1, 2, 3, 12 208. 1931 March 21 Sidney, OH 40.3 84.2 III 1, 8, 14 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-139 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 209. 1931 March 31 Shelby County, OH 40.4 84.1 III 1 210. 1931 June 10 Malinta, OH 41.3 84.0 V 1,500 1, 8, 14 211. 1931 Sept 20 Anna, OH 40.4 84.2 VII 45,400 1, 2, 3, 8,11, 12, 14 212. 1931 Oct 8 Anna, OH 40.4 84.2 III 1, 8, 14 213. 1931 Oct 18 Madison, WI 43.1 89.4 II-III 1 214. 1931 Dec 17 St. Louis, MO 38.6 90.2 II 1 215. 1931 Dec 31 Petersburg, IN 38.5 87.3 NOT RECORDED 1 216. 1934 Nov 12 Rock Island, IL 41.5 90.5 V-VI 5,000 1, 3 217. 1935 Jan 5 Moline, IL 41.5 90.6 III-IV 200 1, 2 218. 1935 Jan 30 Harrison County, MO 40.5 94.0 III 1 219. 1935 Feb 26 Burlington, IA 40.8 91.2 II-III 1 220. 1935 Oct 29 Pike County, IL 39.6 90.8 NOT RECORDED 1 221. 1936 Oct 8 Butler County, OH 39.3 84.4 III 700 1, 8, 14 222. 1936 Dec 25 Cincinnati, OH 39.1 84.5 III 1 223. 1937 March 2 Anna, OH 40.4 84.2 VI-VII 70,000 1, 2, 6, 8, 9, 12, 14 224. 1937 March 3 Anna, OH 40.4 84.2 IV 1, 2, 8, 11, 14 225. 1937 March 3 Anna, OH 40.4 84.2 III 200 1, 8, 14 226. 1937 March 8 Anna, OH 40.4 84.2 VII-VIII 150,000 1, 2, 3, 6, 12, 14 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-140 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 227. 1937 April 23 Anna, OH 40.4 84.2 III 200 1, 8, 14 228. 1937 April 27 Anna, OH 40.4 84.2 III 200 1, 8, 14 229. 1937 May 2 Anna, OH 40.4 84.2 IV 1 230. 1937 June 29 Peoria, IL 40.7 89.6 II 1 231. 1937 Aug 5 Near St. Louis, MO 38.5 90.2 II-III 1 232. 1937 Aug 5 Granite City, IL 38.7 90.2 II 1 233. 1937 Oct 16 Cincinnati, OH 39.1 84.5 II-III 1 234. 1937 Nov 17 Near Centralia, IL 38.6 89.1 V 8,000 1, 2, 3, 6, 12 235. 1938 Feb 12 Porter County, IN 41.6 87.0 IV-V 6,500 1, 2 237. 1939 March 18 Near Jackson Center, OH 40.4 84.0 III 500 1, 8, 14 238. 1939 June 17 Anna, OH 40.4 84.2 IV 400 1, 8, 14 239. 1939 July 9 Anna, OH 40.4 84.2 II 1, 8, 14 240. 1939 Nov 23 Monroe County, IL 38.2 90.1 V-VI 150,000 1, 3 241. 1939 Nov 24 Davenport, IA 41.6 90.6 II-III 1, 2 242. 1940 Jan 8 Louisville, KY 38.3 85.8 II-III 1 243. 1940 May 27 Louisville, KY 38.3 85.8 III 1, 2 244. 1940 May 31 Paducah, KY 37.1 88.6 IV-V 1,000 2 245. 1940 Nov 23 Monroe County, IL 38.2 90.1 VI 150,000 1 246. 1940 Dec 28 Near Evansville, IN 37.9 87.4 III 700 2 247. 1941 Oct 4 St. Louis, MO 38.6 90.2 I 1 248. 1941 Oct 21 Cairo, IL 37.0 89.2 III-IV 1,200 2 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-141 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 249. 1941 Nov 15 Waterloo, IL 38.3 90.2 III 1 250. 1942 Jan Winfield, MO 39.0 90.7 III 1 251. 1942 Jan 14 St. Louis, MO 38.6 90.2 UNKNOWN 600 1 252. 1942 Jan 29 St. Louis, MO 38.6 90.2 UNKNOWN 1 253. 1942 Jan 30 St. Louis, MO 38.6 90.2 UNKNOWN 1 254. 1942 March 1 Kewanee, IL 41.2 89.9 IV-V 3,700 1, 2 255. 1942 March 29 Harrisburg, IL 37.7 88.5 III-IV 200 2 256. 1942 Aug 31 Cairo, IL 37.0 89.2 III-IV 2 257. 1942 Nov 17 East St. Louis, IL 38.6 90.2 III-IV 200 1 258. 1942 Dec 27 Maplewood, MO 38.6 90.3 II 1 259. 1943 April 13 Louisville, KY 38.3 85.8 IV 1 260. 1943 April 18 Waterloo, IL 38.3 90.2 I 1 261. 1943 May 20 West Alton, MO 38.9 90.2 I 1 262. 1943 May 24 West Alton, MO 38.9 90.2 I 1 263. 1943 June 8 Webster Groves, MO 38.6 90.4 III-IV 1 264. 1943 June 15 House Springs, MO 38.4 90.6 I 1 265. 1943 June 18 House Springs, MO 38.4 90.6 I 1 266. 1943 Sept 14 Near St. Louis, MO 38.7 90.3 I 1 267. 1944 Jan 7 Near Jackson, MO 37.5 89.7 III-IV 900 1 268. 1944 March 16 Elgin, IL 42.0 88.3 II 1 269. 1944 Sept 25 St. Louis, MO 38.6 90.2 IV 25,000 1 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-142 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 270. 1944 Nov 13 Anna, OH 40.4 84.2 III-IV 18,000 1, 4, 18 271. 1945 Jan 15 Farmington, MO 37.8 90.4 IV 700 2 272. 1945 March 27 St. Louis, MO 38.6 90.2 II-III 1 273. 1945 May 21 Near St. Louis, MO 38.7 90.2 III-IV 1 274. 1945 Sept 23 Cairo, IL 37.0 89.2 III-IV 2 275. 1945 Nov 13 Cairo, IL 37.0 89.2 III-IV 11,000 2 276. 1946 Feb 24 Centralia, IL 38.5 89.1 IV-V 1,500 1, 2 277. 1946 Oct 7 Near Chloride, MO 37.5 90.6 IV-V 32,000 2 278. 1946 Nov 7 Washington County, MO 38.0 90.7 II-III 1 279. 1947 March 16 Kane County, IL 42.1 88.3 IV 1 280. 1947 May 6 Milwaukee, WI 43.0 87.9 V 3,000 1, 2 281. 1947 June 29 Near St. Louis, MO 38.4 90.2 VI 15,000 1, 3 282. 1947 Aug 9 Branch County, MI 42.0 85.0 VI 70,000 1, 2, 3 283. 1948 Jan 5 Centralia, IL 38.5 89.1 IV-V 300 1, 13 284. 1948 Jan 15 Madison County, WI 43.2 89.7 IV-V 1 285. 1948 April 20 Iowa City, IA 41.7 91.5 III-IV 1 286. 1949 June 8 Ste. Genevieve, MO 38.0 90.1 III 300 1 287. 1949 Aug 11 Clayton, MO 38.7 90.3 II 1 288. 1949 Aug 26 Defiance, MO 38.6 90.8 II-III 1 289. 1950 Feb 8 Lebanon, MO 37.7 92.7 V 1 290. 1950 April 20 Dayton, OH 39.8 84.2 IV 1, 8, 14 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-143 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 291. 1951 Sept 19 Near Florissant, MO 38.9 90.2 III-IV 1,200 1 292. 1952 Jan 7 Champaign County, IL 40.3 88.3 II-III 1 293. 1953 May 6 Cairo, IL 37.0 89.2 III 2 294. 1953 May 15 Cairo, IL 37.0 89.2 III 2 295. 1953 Sept 11 Near Roxana, IL 38.6 90.1 VI 6,000 1, 3 296. 1953 Dec 30 Centralia, IL 38.5 89.1 IV 1,200 1 297. 1954 Aug 9 Petersburg, IN 39.5 87.3 IV-V 1, 2 298. 1955 April 9 Near Sparta, IL 38.1 89.8 VI 20,000 1, 3 299. 1955 May 29 Ewing, IL 38.1 88.9 III-IV 1 300. 1956 Jan 27 Anna, OH 40.4 84.2 V 2,000 1, 2, 8, 14 301. 1956 March 13 Fulton County, IL 40.5 90.2 IV 2,000 1 302. 1956 July 18 Oostburg, WI 43.6 87.8 IV 1 303. 1956 Oct 13 Near Milwaukee, WI 42.8 87.9 IV 1 304. 1956 Nov 25 Wayne County, MO 37.1 90.6 VI 21,500 2, 3 305. 1957 Jan 8 Waupun, WI 43.6 88.7 III-IV 1 306. 1957 March 26 Paducah, KY 37.1 88.6 V 300 2, 3 307. 1958 Jan 27 Ballard County, KY 37.0 89.0 V 300 3 308. 1958 Nov 7 Wabash County, IL 38.4 87.9 VI 33,300 1, 2, 3, 9 310. 1962 June 26 Saline County, IL 37.7 88.5 V-VI 17,500 2, 3 311. 1963 Aug 2 McCracken County, KY 37.0 88.8 IV-V 2,600 2, 3 312. 1965 March 6 Crawford County, MO 37.8 91.2 VI-VII 2 CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-144 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 313. 1965 Aug 14 Pulaski County, IL 37.1 89.2 VI-VII 400 2, 3 314. 1965 Aug 15 Cape Girardeau County, MO 37.4 89.5 VI 3 315. 1965 Oct 20 Washington County, MO 37.8 91.1 VI 160,000* 1, 2, 3 316. 1967 Feb 2 Lansing, MI 42.7 84.5 IV 1 317. 1967 July 21 Madison County, MO 37.5 90.4 VI 2, 3 318. 1967 Aug 5 Jefferson County, MO 38.3 90.6 II 1 319. 1968 Nov 9 Hamilton County, IL 38.0 88.5 VII 580,000 1, 2, 3 320. 1968 Dec 11 Louisville, KY 38.3 85.8 V 1 321. 1969 Jan 20 Farmington, MO 37.8 90.4 III 1 322. 1971 Feb 12 Wabash County, IL 38.5 87.9 IV 1,300 1 323. 1972 June 9 St. Francois County, MO 37.7 90.4 IV 1 324. 1972 June 19 Ballard County, KY 37.0 89.1 IV 1 325. 1972 Sept 15 Lee County, IL 41.6 89.4 VI 40,000 1, 5 326. 1973 Jan 7 Hopkins County, KY 37.4 87.3 III 1 327. 1973 Jan 12 St. Francois County, MO 37.9 90.5 III 1 328. 1973 April 18 St. Clair County, IL 38.5 90.2 II-III 1 329. 1974 March 27 St. Louis, MO 38.5 90.1 II-III 17 330. 1974 April 3 Southern Illinois 38.6 88.1 VI 17 331. 1974 April 5 Eastern Missouri 38.6 90.9 IV or less 17 332. 1974 June 5 Kentucky 38.6 84.8 NOT RECORDED 17

• 245,000 mi 2 according to Reference 2.

CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-145 REV. 11, JANUARY 2005 DATE LOCATION NORTH LATITUDE WEST LONGITUDE MAXIMUM INTENSITY (MM) FELT AREA (mi²) REFERENCES 333. 1974 June 5 Southern Illinois 38.6 89.9 V 17 334. 1974 Aug 22 Southern Illinois 38.2 89.7 V 17 335. 1976 April 8 Stinesville, IN 39.3 86.8 V 18 NOTE Blank spaces in Table 2.5-4 indicate that data is not available.

CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-146 REV. 11, JANUARY 2005 (References for Table 2.5-4)

1. M.M. Varma, "Seismicity of the Eastern Half of the United States," unpublished Ph.D. dissertation, Indiana University, Bloomington, 176 pp., 1975. 2. J. Docekal, "Earthquakes of the Stable Interior with Emphasis on the Midcontinent," Ph.D. dissertation, University of Nebraska, Lincoln, Vols. 1 and 2, 1970. 3. J.L. Coffman and C.A. von Hake (eds.), "Earthquake History of the United States," U.S. Dept. of Commerce, NOAA, Environmental Data Service, Boulder, Colo., Publication 41-1, revised edition through 1970, 1973. 4. B.C. Moneymaker, "Some Earthquakes in Tennessee and Adjacent States (1699 to 1850)," Tenn. Acad. Sci. Jour. 29, 3:224-233, 1954. 5. P.C. Heigold, "Notes on the Earthquake of September 15, 1972 in Northern Illinois," Illinois State Geol. Surv. Environmental Geol. Notes, No. 59, 1972. 6. R.R. Heinrich, "A Contribution to the Seismic History of Missouri," Seismol. Soc. Amer Bull. 31, 3:187-224, 1941. 7. B.C. Moneymaker, Tennessee Valley Authority, unpublished report, 1964. 8. E.A. Bradley and T.J. Bennett, "Earthquake History of Ohio," Seismol. Soc. Amer. Bull. 55, 4:745-752, 1965. 9. O.W. Nuttli, "State-of-the-Art for Assessing Earthquake Hazards in the United States," U.S. Army Engineers Waterways Experiment Station, Report 1, Design Earthquakes for the Central United States, 1973. 10. B.C. Moneymaker, "Earthquakes in Tennessee and Nearby Sections of Neighboring States (1901-1925)," Tenn-Acad. Sci. Jour. 32, 2:91-105, 1957. 11. R.R. Heinrich, "The Mississippi Valley Earthquake of June 20, 1947," Seismol. Soc. Amer. Bull. 40:7-19, 1950. 12. U.S. Coast and Geodetic Survey, "United States Earthquakes, 1920-1935 and United States Earthquakes, 1936-1940, "U.S. Department of Commerce, Environmental Science Services Administration, National Earthquake Information Center, 1968 reissue and 1969 reissue, respectively. 13. B.C. Moneymaker, "Earthquakes in Tennessee and Nearby Sections of Neighboring States (1926-1950)," Tenn. Acad. Sci. Jour. 33, 3:224-239, 1958. 14. E.F. Pawlowicz, "Earthquake Statistics for Ohio," Ohio Jour. Sci. 75, 2:111, March 1975. 15. B.C. Moneymaker, "Earthquakes of Kentucky," unpublished, Tenn. Valley Authority, undated.

CPS/USAR TABLE 2.5-4 (Cont'd) CHAPTER 02 2.5-147 REV. 11, JANUARY 2005 16. C.G. Rockwood, "Notes on American Earthquakes," Amer. Jour. Sci., Vol. 21, No. 13, 3rd series, 1984. 17. U.S. Geological Survey, "Preliminary Determination of Epicenters," USGS monthly publication. 18. U.S. Geological Survey, Earthquake Information, Bulletin, USGS, bimonthly publication. 19. M. C. Hansen "Earthquakes in Ohio," Ohio Division Geological Survey, Education Leaflet No. 9 (1975).

CPS/USAR CHAPTER 02 2.5-148 REV. 11, JANUARY 2005 TABLE 2.5-5 EARTHQUAKES OCCURRING OVER 200 MILES FROM THE SITE FELT AT THE CLINTON SITE EPICENTER LOCATION (degrees)

DATE MAXIMUM MODIFIED MERCALLI INTENSITY LOCALITY N. LAT. W. LONG. FELT AREA (mi²)DISTANCE FROM SITE (mi)1811 December 16 XI Northeastern Arkansas Gulf Coast Tectonic

Province 35.5 90.5 2,000,000 3501812 X-XI January 23 New Madrid, Missouri Gulf Coast Tectonic

Province 36.6 89.5 2,000,000 2851812 XI-XII February 7 New Madrid, Missouri Gulf Coast Tectonic

Province 36.6 89.5 2,000,000 2851886 X August 31 Charleston, South Carolina Atlantic Coast Tectonic Province 32.9 80.0 2,000,000 7001895 VIII October 31 Charleston, Missouri Gulf Coast Tectonic Province 37.0 89.4 1,000,000 2151937 VII-VIII March 8 Anna, Ohio Central Stable Region 40.4 84.2 150,000 245 CPS/USAR CHAPTER 02 2.5-149 REV. 11, JANUARY 2005 TABLE 2.5-6 DIRECT SHEAR TEST DATA D BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE* GEOLOGIC UNIT DRY DENSITY (lb/ft³)

FIELD MOISTURE CONTENT (percent) NORMAL PRESSURE (lb/ft²)

YIELD STRENGTH (lb/ft²)

PEAK STRENGTH (lb/ft²) D-10 650.0 SM Salt Creek Alluvium 107 17.3 1000 483 725 D-10 650.0 SM Salt Creek Alluvium 107 17.3 2000 705 1060 D-10 647.0 SM.SP Salt Creek Alluvium 85 34.1 3000 960 1440 D-10 647.0 SM.SP Salt Creek Alluvium 85 34.1 4000 985 1480 D-10 642.0 SM.GP Salt Creek Alluvium 131 8.3 3000 1025 1540 D-10 642.0 SM.GP Salt Creek Alluvium 131 8.3 6000 985 1480 D-10 638.0 ML Illinoian Glacial Till 140 8.8 2000 2880 4320 D-10 587.0 ML Illinoian Glacial Till 141 6.9 6000 4334 6500Limit D-11 619.8 ML Illinoian Glacial Till 144 6.2 4000 --- 6500Limit D-11 559.8 ML.CL Illinoian Glacial Till 134 9.2 6000 --- 6500Limit D-11 524.8 ML Illinoian Glacial Till 125 12.6 6000 3770 5640 D-24 649.0 SP Salt Creek Alluvium 106 18.1 1000 420 630 D-24 649.0 SP Salt Creek Alluvium 106 18.1 2000 1035 1550 D-24 646.0 SP Salt Creek Alluvium 100 21.7 3000 380 570 D-24 646.0 SP Salt Creek Alluvium 100 21.7 4000 486 730 D-24 636.0 SP Salt Creek Alluvium 109 17.5 3000 466 700 D-24 636.0 SP Salt Creek Alluvium 109 17.5 6000 900 1350 D-30 645.9 CL Illinoian Glacial Till 138 7.4 2500 4000 6000 D-34 619.8 ML Illinoian Glacial Till 142 7.3 5000 --- 6500 Limit __________________________

• See Figure 2.5-355 for definition of soil type symbols.

CPS/USAR CHAPTER 02 2.5-150 REV. 11, JANUARY 2005 TABLE 2.5-7 DIRECT SHEAR TEST DATA B BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL* TYPE GEOLOGIC UNIT FIELD DRY DENSITY (lb/ft³)

FIELD MOISTURE CONTENT (percent) NORMAL PRESSURE (lb/ft²)

YIELD STRENGTH (lb/ft²)

PEAK STRENGTH (lb/ft²) B-5 667.7 SM Wisconsinan Glacial Till 131 10.1 2000 906 1360 B-5 667.7 SM Wisconsinan Glacial Till 131 10.1 4000 1882 2830 B-8 661.8 SM Salt Creek Alluvium 118 12.3 2000 1465 2200 B-8 661.8 SW Salt Creek Alluvium 118 12.3 4000 2465 3700

______________________________

• See Figure 2.5-355 for definition of soil type symbols.

CPS/USAR CHAPTER 02 2.5-151 REV. 11, JANUARY 2005 TABLE 2.5-8 UNCONFINED COMPRESSION TEST DATA REMOLDED SAMPLES**

REMOLDED DATA BORING NUMBER ELEVATION (feet) SOIL TYPE* GEOLOGIC UNIT DRY DENSITY (lb/ft³) COMPACTION EFFORT* (percent) MOISTURE CONTENT (percent) SHEAR STRENGTH (lb/ft²) S-5 714.0 CL Wisconsinan Glacial Till 113 89 16.3 1542 S-5 714.0 CL Wisconsinan Glacial Till 112 88 16.2 1420 S-5 714.0 CL Wisconsinan Glacial Till 109 86 17.9 850 S-5 714.0 CL Wisconsinan Glacial Till 108 85 18.3 800

__________________________

• Modified AASHO (T-180). ** As compacted.

CPS/USAR CHAPTER 02 2.5-152 REV. 11, JANUARY 2005 Table 2.5-9 UNCONFINED COMPRESSION TEST DATA REMOLDED SAMPLES AS COMPACTED REMOLDED DATA BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE* GEOLOGIC UNIT DRY DENSITY (lb/ft³) COMPACTION EFFORT**

(percent) MOISTURE CONTENT (percent) SHEAR STRENGTH (lb/ft²) S-9 712.2 -

707.2 CL Wisconsinan Glacial Till 122 102.1 13.4 2,960 S-9 712.2 -

707.2 CL Wisconsinan Glacial Till 122 102.1 13.1 3,500 S-9 712.2 -

707.2 CL Wisconsinan Glacial Till 120 100.4 13.6 700 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 118 91.0 15.2 690 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 122 94.2 14.2 1,020 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 119 91.9 13.7 950 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 124 95.8 11.6 2,390 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 125 96.5 12.1 2,880 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 120 92.6 13.9 870 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 116 89.6 15.3 720 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 117 90.3 15.5 520 S-10 702.6 -

697.6 CL Wisconsinan Glacial Till 121 93.4 12.4 1,760 S-12 719.2 -

712.2 CL Wisconsinan Glacial Till 117 95.9 14.8 1,560 S-12 719.2 -

712.2 CL Wisconsinan Glacial Till 105 86.0 14.7 1,440 ____________________________

• See Figure 2.5-355 for the definition of soil type symbols. ** AASHO Test Designation T-180.

CPS/USAR TABLE 2.5-9 (Cont'd) CHAPTER 02 2.5-153 REV. 11, JANUARY 2005 REMOLDED DATA BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE* GEOLOGIC UNIT DRY DENSITY (lb/ft³) COMPACTION EFFORT**

(percent) MOISTURE CONTENT (percent) SHEAR STRENGTH (lb/ft²) S-12 719.2 -

712.2 CL Wisconsinan Glacial Till 116 95.0 15.5 1,400 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 125 96.1 11.6 1,380 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 118 90.7 14.2 585 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 113 86.9 16.1 305 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 114 87.7 12.7 1,320 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 113 86.9 15.2 475 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 121 93.1 13.1 1,220 S-14 705.2 -

700.2 CL Wisconsinan Glacial Till 115 88.4 15.1 460 S-14 727.2 -

720.2 CL Wisconsinan Glacial Till 103 86.6 14.5 1,520 S-14 727.2 -

720.2 CL Wisconsinan Glacial Till 115 96.6 12.9 4,000

____________________________

• See Figure 2.5-355 for the definition of soil type symbols. ** AASHO Test Designation T-180.

CPS/USAR CHAPTER 02 2.5-154 REV. 11, JANUARY 2005 TABLE 2.5-10 CONSOLIDATED-UNDRAINED TRIAXIAL TEST DATA WITH PORE PRESSURE MEASUREMENTS STRESSES at (1'-3') /2 maximum BORING NUMBER ELEVATION (ft-in.) GEOLOGIC UNIT CONSOLIDATION PRESSURE c' = 3' (lb/ft2) PEAK SHEAR STRENGTH (1'-3') /2 maximum (lb/ft2) u 1' 3' (1'+3')/2(lb/ft2) D-8 631.7 Illinoian Glacial Till 3,000 19,592 -418 52,676 13,492 33,084 D-8 591.7 Illinoian Glacial Till 6,000 14,955 -331 43,446 13,536 28,491 D-48 709.3 Wisconsinan Glacial Till 2,000 1,009 187 3,831 1,813 2,822 D-48 704.3 Wisconsinan Glacial Till 4,000 6,434 -1,584 18,451 5,584 12,018 D-48 689.3 Wisconsinan Glacial Till 6,000 7,089 -158 20,335 6,158 13,247

_________________________

CPS/USAR CHAPTER 02 2.5-155 REV. 11, JANUARY 2005 TABLE 2.5-11 CONSOLIDATED-UNDRAINED TRIAXIAL COMPRESSION TEST DATA WITH PORE PRESSURE MEASUREMENTS PEAK SHEAR CONSOLIDATION PRESSURE STRENGTH STRESSES AT c' = 3' 1'-3' MAX./2 1'-3'/2 MAX (lb/ft2) 1'+3'/2 BORING NUMBER ELEVATION

(ft-in.) GEOLOGIC UNIT (lb/ft ²)

(lb/ft ²)

u 1' 3' (lb/ft ²) P-38 648.5 Illinoian Glacial Till 6,480 19,646

• 45,7736,48026,127 P-38 572.9 Lacustrine Deposits 10,000 7,415 4,29120,5315,70213,116 H-23 707.3 Wisconsinan Glacial Till 2,160 3,414 -4039,3912,5635,977 H-23 692.3 Interglacial Soil 11,808 7,672 5,12622,0246,68214,353 H-38 673.4 Illinoian Glacial Till 3,528 11,245 -50426,5224,03215,277 H-3 645.1 Illinoian Glacial Till 2,016 4,611 -32911,5672,3456,956 H-20 721.8 Wisconinan Glacial Till 1,440 9,052 -4,46424,0095,90414,956

____________________________

• Test performed on specimen at field moisture content; pore pressure measured negligible.

CPS/USAR TABLE 2.5-11 (Cont'd) CHAPTER 02 2.5-156 REV. 11, JANUARY 2005 CONSOLIDATION PEAK SHEAR PRESSURE STRENGTH STRESSES AT c' = 3' 1'-3' MAX./2 1'-3'/2 MAX (lb/ft2) 1'+ 3'/2 BORING NUMBER ELEVATION

(ft-in.) GEOLOGIC

UNIT (lb/ft ²)

(lb/ft²) u 1' 3' (lb/ft ²)

H-24 670.7 Salt Creek Alluvium 388 1,610 -1,2244,8321,6123,222 H-38 712.9 Wisconinan Glacial Till 4,997 2,813 -24510,8685,2428,055 H-38 687.9 Interglacial Soil 4,608 3,888 85011,5343,7587,646 H-25 633.7 Illinoian Glacial Till 8,640 26,403 -3,31264,75811,95238,355 H-13 673.6 Salt Creek Alluvium 389 1,785 -5354,4949242,709 H-6 504.3 Illinoian Glacial Till 9,994 7,667 3,98921,3386,00513,672 H-25 674.7 Salt Creek Alluvium 100 509 -1151,235215725 CPS/USAR CHAPTER 02 2.5-157 REV. 11, JANUARY 2005 TABLE 2.5-12 UNCONSOLIDATED-UNDRAINED TRIAXIAL COMPRESSURE TEST DATA REMOLDED SAMPLES

• REMOLDED DATA BORING NUMBER ELEVATION

(ft-in.)

SOIL TYPE** GEOLOGIC

UNIT DRY DENSITY (lb/ft³) COMPACTION EFFORT***

(percent) MOISTURE CONTENT (percent) CONFINING PRESSURE

(lb/ft²) SHEAR STRENGTH

(lb/ft²) S-14 705.2 to 700.2 CL Wisconsinan Glacial Till 116 89.2 14.9 2000 276 S-14 705.2 700.2 CL Wisconsinan Glacial Till 119 91.5 13.7 4000 877 S-14 705.2 to 700.2 CL Wisconsinan Glacial Till 116 89.2 14.9 6000 544

________________________

• As compacted. ** See Figure 2.5-355 for definition of soil type symbols.

• • • AASHO Test Designation T-180.

CPS/USAR TABLE 2.5-12 (Cont'd) CHAPTER 02 2.5-158 REV. 11, JANUARY 2005 REMOLDED DATA BORING NUMBER ELEVATION

(feet) SOIL TYPE** GEOLOGIC

UNIT DRY DENSITY (lb/ft³) COMPACTION EFFORT***

(percent) MOISTURE CONTENT (percent) CONFINING PRESSURE

(lb/ft²) SHEAR STRENGTH

(lb/ft²)

Combined sample from:

P-30 P-32 P-33 P-35 727-717 CL Wisconsinan Glacial Till 130.2 96.8 10.1 1500 4273 Combined sample from:

P-30 P-33 P-35 713-698 CL Wisconsinan Glacial Till 126.7 97.2 11.0 1500 2424 Combined sample from:

P-37 P-39 P-42 730-720 CL Wisconsinan Glacial Till 123.0 91.8 11.7 1500 1106 Combined sample from:

P-30 P-32 to P-35 P-37 P-39 P-42 727-692 CL Wisconsinan Glacial Till 129.2 -- 9.9 1500 3674

CPS/USAR TABLE 2.5-12 (Cont'd) CHAPTER 02 2.5-159 REV. 11, JANUARY 2005 REMOLDED DATA BORING NUMBER ELEVATION

(feet) SOIL TYPE** GEOLOGIC

UNIT DRY DENSITY (lb/ft³) COMPACTION EFFORT***

(percent) MOISTURE CONTENT (percent) CONFINING PRESSURE

(lb/ft²) SHEAR STRENGTH

(lb/ft²)

Combined sample from:

P-30 P-33 P-35 713 to 698 CL Wisconsinan Glacial Till 122.9 91.0 8.7 1728 1364 (Saturated)

Combined sample from:

P-34 P-37 P-39 P-42 720 to 700 CL Wisconsinan Glacial Till 129.8 97.0 10.3 1728 1631 (Saturated)

CPS/USAR CHAPTER 02 2.5-160 REV. 11, JANUARY 2005 TABLE 2.5-13 CONSOLIDATED-UNDRAINED TRIAXIAL TEST DATA WITH PORE PRESSURE MEASUREMENTS REMOLDED SAMPLES*

REMOLDED DATA PEAK SHEAR STRESSES AT STRENGTH BORING NUMBER ELEVATION (ft-in.)

GEOLOGIC UNIT MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft3) COMPACTION EFFORT** (percent) CONSOLIDATION PRESSURE c'=3' (lb/ft²) 1'-3'/2 Max (lb/ft2) 1'-3'/2 Max (lb/ft2) u 1' 3' 1'+3'/2 (lb/ft

2) S-10 717.6- Wisconsinan 15.5 121 89.4 2,620 3,069 -43 8,801 2,663 5,732 712.6 Glacial Till S-10 717.6- Wisconsinan 13.4 119 87.9 4,075 3,504 1,396 9,688 2,679 6,184 712.6 Glacial Till S-10 717.6- Wisconsinan 12.6 123 90.8 7,257 5,942 1,958 17,183 5,299 11,241 712.6 Glacial Till S-10 702.6- Wisconsinan 14.0 118 91.1 2,000 1,801 518 5,014 1,412 3,212 697.6 Glacial Till S-10 702.6- Wisconsinan 13.5 120 92.6 4,000 2,170 2,260 6,069 1,728 3,899 697.6 Glacial Till S-10 702.6- Wisconsinan 13.8 121 93.4 6,000 4,523 10,166 13,280 4,234 8,757 697.6 Glacial Till ________________________

• As compacted. ** AASHO Test Designation T-180.

CPS/USAR CHAPTER 02 2.5-161 REV. 11, JANUARY 2005 TABLE 2.5-14 CONSOLIDATED-UNDRAINED TRIAXIAL TEST DATA WITH PORE PRESSURE MEASUREMENTS REMOLDED SAMPLES REMOLDED DATA PEAK SHEAR STRESSES AT STRENGTH BORING NUMBER ELEVATION (ft-in.)

GEOLOGIC UNIT MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft3) COMPACTION EFFORT** (percent) CONSOLIDATION PRESSURE c'=3' (lb/ft²) 1'-3'/2 Max (lb/ft2) 1'-3'/2 Max (lb/ft2) u 1' 3' 1'+3'/2 (lb/ft

2) Combined sample from: P-30 713-6 Wisconsinan 8.5 123.6 94.8 2,000 2,983 +173 7,793 1,827 4,810 P-33 98 Glacial P-35 Till Combined sample from: P-37 730-7 Wisconsinan 12.8 123.1 91.6 8,000 4,909 +3,960 13,858 4,040 8,949 P-39 20 Glacial P-42 Till CPS/USAR CHAPTER 02 2.5-162 REV. 11, JANUARY 2005 TABLE 2.5-15 CONSOLIDATED-UNDRAINED TRIAXIAL TEST DATA WIITH PORE PRESSURE MEASUREMENT REMOLDED SAMPLES PEAK SHEAR STRESSES AT REMOLDED DATA STRENGTH BORING NUMBER ELEVATION (ft-in.)

GEOLOGIC UNIT MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft3) COMPACTION EFFORT** (percent) CONSOLI8DATION PRESSURE c'=3' (lb/ft²) 1'-3'/2 Max (lb/ft2) 1'-3'/2 Max (lb/ft2) u 1' 3' 1'+3'/2 (lb/ft

2) Structural Salt Creek 12.1a 118.8 a 67.0 a 1,008 7,235 -2,995 18,473 4,003 11,238 fill borrow: Alluvium Combined bulk

- a - a - a 4,032 10,125 -2,174 26,456 6,003 16,331 sample from: Salt Creek G-18 663-654 Alluvium 12.4b 118.0 b 65.0 b 4,032 11,519 -2,059 29,128 6,091 17,610 G-19 673-663 G-20 657-647

- b - b - b 7,000 16,016 -2,505 41,537 9,505 25,521 - b - b - b 9,936 18,238 -1,022 47,433 10,958 29,196 Structural Salt Creek 9.0c 120.2 c 70.5 c 876 10,832 -5,501 28,043 6,380 17,212 fill borrow: Alluvium Combined bulk

- c - c - c 3,975 11,655 -3,370 30,678 7,368 19,023 sample from: G-18 663-654

- c - c - c 6,984 15,290 -2,808 40,375 9,794 25,084 G-19 673-663 G-20 657-647 Salt Creek 8.4d 129.4 d 91.0 d 3,984 19,104 -5,501 47,698 9,490 28,594 Alluvium

- d - d - d 5,688 19,102 -4,090 47,983 9,778 28,880 - d - d - d 8,640 18,030 -3,470 48,170 12,110 30,140 CPS/USAR TABLE 2.5-15 (Cont'd) CHAPTER 02 2.5-163 REV. 11, JANUARY 2005 ______________________

• ASTM D2049-69 Relative Density of Cohesionless Soils.

a M/TX/CU/pp: Multiphase triaxial compression test, saturated, consolidated-undrained with pore pressure measurement; utilizing the same specimen initially molded at 12.1% moisture content and 118.8 lb/ft 3 dry density.

b M/TX/CU/pp, utilizing the same specimen initially molded at 12.4% moisture content and 118.0 lb/ft 3 dry density.

c M/TX/CU/pp, utilizing the same specimen initially molded at 9.0% moisture content and 120.2 lb/ft 3 dry density.

d M/TX/CU/pp, utilizing the same specimen initially molded at 8.4% moisture content and 129.4 lb/ft 3 dry density.

CPS/USAR CHAPTER 02 2.5-164 REV. 11, JANUARY 2005 TABLE 2.5-16 MOISTURE AND DENSITY DATA ULTIMATE HEAT SINK BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT DRY DENSITY (lb/ft³) MOISTURE CONTENT (%) H-1 667.7 ML Salt Creek Alluvium 93 29.2 H-6 619.3 ML Illinoian Glacial Till 136 8.6 H-6 579.3 CL Illinoian Glacial Till 129 10.3 H-6 425.8 CH Pre-Illinoian Lacustrine Deposit 120 15.2 H-14 635.3 ML Illinoian Glacial Till 140 9.5 H-15 691.3 CL Interglacial Zone 105 17.9* H-17 669.6 ML Salt Creek Alluvium 98 24.1 H-17 641.6 ML Illinoian Glacial Till 139 8.7 H-20 731.8 ML Wisconsinan Glacial Till 119 13.4 H-20 721.8 ML Wisconsinan Glacial Till 130 9.5 H-20 716.8 ML Wisconsinan Glacial Till 134 10.3* H-20 706.8 ML Wisconsinan Glacial Till 136 8.6 H-20 691.8 ML Interglacial Soil 113 16.6 H-20 672.3 ML Illinoian Glacial Till 135 10.5 H-22 540.8 ML Pre-Illinoian Glacial Till 128 12.3 H-23 730.8 ML Weathered Loess109 12.8 H-23 707.3 ML Wisconsinan Glacial Till 121 13.6 _________________

• Saturated CPS/USAR TABLE 2.5-16 (Cont'd) CHAPTER 02 2.5-165 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT DRY DENSITY (lb/ft³) MOISTURE CONTENT (%) H-23 692.3 ML Interglacial Soil 114 17.4 H-23 677.8 ML Illinoian Glacial Till 138 9.1 H-30 638.5 ML Illinoian Glacial Till 145 6.8 H-30 633.5 ML Illinoian Glacial Till 141 7.4 H-31 667.1 ML Illinoian Glacial Till 139 8.7 H-32 616.1 ML Illinoian Glacial Till 134 8.4 H-32 575.1 CL Lacustrine Deposit 116 14.6 H-36 622.7 ML Illinoian Glacial Till 137 9.4 H-36 597.2 ML Illinoian Glacial Till 139 7.7 H-37 689.9 ML Interglacial Soil 122 6.5 H-38 712.9 ML Wisconsinan Glacial Till 124 10.3 H-38 707.9 ML Wisconsinan Glacial Till 122 12.4 H-38 687.9 ML Inter-Glacial Soil 113 17.4 H-38 673.4 ML Illinoian Glacial Till 135 8.6 H-40 672.6 ML Salt Creek Alluvium 86 24.7

CPS/USAR CHAPTER 02 2.5-166 REV. 11, JANUARY 2005 TABLE 2.5-17 UNCONSOLIDATED-UNDRAINED TRIAXIAL TEST DATA ULTIMATE HEAT SINK BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT CONFINING PRESSURE (lb/ft²) SHEAR STRENGTH (lb/ft²)

DRY DENSITY (lb/ft³) MOISTURE CONTENT (%) H-14 645.3 ML Illinoian Glacial Till 1900 6902* 143 8.2* H-16 652.8 ML Illinoian Glacial Till 2002 1056 137 8.3 H-20 691.8 ML Interglacial Soil 3197 3736* 115 17.9* H-23 730.8 ML Weathered Loess 346 610* 114 20.4* H-23 697.3 ML Interglacial Soil 2534 2974* 100 25.0*

H-30 658.0 ML Illinoian Glacial Till 2693 7569 135 7.9 H-30 618.5 ML Illinoian Glacial Till 5396 954 120 11.1 H-38 707.9 ML Wisconsinan Glacial Till 3090 1238* 120 14.1* H-38 692.9 CL Interglacial Soil 2189 1626* 104 21.4*

H-23 662.8 ML Illinoian Glacial Till 4750 1400 128 13.4 H-6 564.3 ML Illinoian Glacial Till 7200 6840* 129 10.7*

__________________________

• Saturated CPS/USAR TABLE 2.5-17 (Cont'd) CHAPTER 02 2.5-167 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT CONFINING PRESSURE (lb/ft²) SHEAR STRENGTH (lb/ft²)

DRY DENSITY (lb/ft³) MOISTURE CONTENT (%) H-17 672.6 ML Salt Creek Alluvium 288 850* 92 24.7* H-14 668.3 SM Salt Creek Alluvium 360 916* 108 19.1* H-9 692.1 ML Interglacial Soil 2600 1735* 108 20.5* H-15 651.8 ML Illinoian Glacial Till 4565 7745* 138 8.5* H-22 554.8 ML Pre-Illinoian Glacial Till 6984 2580* 127 13.0* H-19 675.0 ML Salt Creek Alluvium 200 396* 110 22.1* H-15 696.3 ML Interglacial Soil 4850 1350* 82.4 35.4*

__________________________

• Saturated CPS/USAR CHAPTER 02 2.5-168 REV. 11, JANUARY 2005 TABLE 2.5-18 DYNAMIC TRIAXIAL COMPRESSION TEST DATA STATION SITE BORINGS BORING NUMBER ELEVATION (ft-in.) SOIL TYPE FIELD MOISTURE CONTENT (percent) DRY DENSITY (lb/ft³) SINGLE AMPLITUDE SHEAR STRAIN (percent)

MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) P-14 714.3 ML 16.6 117 0.0141 1.25 x 10 6 17 0.0264 1.07 x 10 6 16 0.0521 0.85 x 10 6 16 (Wisconsinan Glacial Till) 0.1268 0.55 x 10 6 18 0.2508 0.40 x 10 6 19 0.5191 0.25 x 10 6 20 1.2540 0.11 x 10 6 20 1.9153 0.09 x 10 6 21 2.5205 0.08 x 10 6 21 P-14 624.3 CL 8.3 139 0.0127 2.60 x 10 6 11 0.0262 1.99 x 10 6 12 (Illinoian Glacial Till) 0.0508 1.53 x 10 6 15 0.1272 0.91 x 10 6 19 0.2502 0.66 x 10 6 21 0.5077 0.46 x 10 6 22 0.9921 0.35 x 10 6 21 1.9841 0.24 x 10 6 19 3.8697 0.33 x 10 6 20 P-15 692.3 ML 18.0 111 0.0144 1.18 x 10 6 14 0.0277 0.97 x 10 6 12 (Interglacial Zone) 0.0531 0.84 x 10 6 11 0.1037 0.61 x 10 6 12 0.2529 0.39 x 10 6 15 0.5059 0.26 x 10 6 18 1.2647 0.15 x 10 6 21 2.5170 0.09 x 10 6 25 3.6941 0.06 x 10 6 29 5.0841 0.05 x 10 6 32 CPS/USAR CHAPTER 02 2.5-169 REV. 11, JANUARY 2005 TABLE 2.5-19 DYNAMIC TRIAXIAL COMPRESSION TEST DATA ADDITIONAL STATION SITE BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) P-32 657.4 ML 7.9 135 0.0063 2.45 x 10 6 11 (Illinoian Glacial Till) 0.0117 2.09 x 10 6 13 0.0175 1.78 x 10 6 14 0.0400 1.14 x 10 6 17 0.0751 0.78 x 10 6 21 0.1562 0.43 x 10 6 20 0.2674 0.26 x 10 6 19 0.5474 0.14 x 10 6 18 P-32 617.4 ML 9.6 133 0.0065 1.05 x 10 6 18 (Illinoian Glacial Till) 0.0167 0.76 x 10 6 15 0.0283 0.59 x 10 6 17 0.0575 0.40 x 10 6 18 0.1133 0.27 x 10 6 18 0.2196 0.17 x 10 6 21 0.3798 0.12 x 10 6 20 0.7596 0.08 x 10 6 20 1.5248 0.05 x 10 6 19 P-32 547.4 ML 10.5 131 0.0242 2.15 x 10 6 (Pre-Illinoian Glacial Till) 0.0653 1.21 x 10 6 0.1306 0.82 x 10 6 19 0.2177 0.62 x 10 6 20 0.3508 0.47 x 10 6 20 0.9193 0.27 x 10 6 21 1.8870 0.18 x 10 6 21 P-36 668.2 ML 13.0 132 0.0670 0.14 x 10 6 28 (Illinoian Glacial Till, low blow count 0.0131 0.11 x 10 6 26 material and test sample may also have 0.0304 0.07 x 10 6 24 been slightly disturbed) 0.0628 0.05 x 10 6 24 0.1195 0.03 x 10 6 25 0.2403 0.02 x 10 6 21 0.4006 0.02 x 10 6 20

CPS/USAR TABLE 2.5-19 (Cont'd) CHAPTER 02 2.5-170 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) P-36 578.7 ML 8.5 138 0.0059 1.98 x 10 6 13 (Illinoian Glacial Till) 0.0098 1.80 x 10 6 16 0.0264 1.13 x 10 6 18 0.0540 0.79 x 10 6 19 0.1083 0.55 x 10 6 20 0.2210 0.39 x 10 6 19 0.3657 0.30 x 10 6 18 0.7353 0.23 x 10 6 15 P-36 519.7 CL 14.5 122 0.0095 1.97 x 10 6 12 (Pre-Illinoian Glacial Till) 0.0348 1.40 x 10 6 12 0.0667 1.13 x 10 6 14 0.1275 0.86 x 10 6 15 0.2633 0.63 x 10 6 16 0.4468 0.50 x 10 6 16 0.9033 0.36 x 10 6 16 P-38 633.9 ML 8.0 137 0.0055 2.54 x 10 6 15 (Pre-Illinoian Glacial Till) 0.0148 1.89 x 10 6 14 0.0272 1.41 x 10 6 15 0.0568 1.00 x 10 6 18 0.1069 0.72 x 10 6 19 0.1755 0.55 x 10 6 20 0.7887 0.23 x 10 6 1.5380 0.17 x 10 6 18 P-38 588.9 ML 9.2 134 0.0092 0.67 x 10 6 18 (Illinoian Glacial Till) 0.0160 0.54 x 10 6 17 0.0314 0.38 x 10 6 19 0.0624 0.27 x 10 6 18 0.0936 0.22 x 10 6 19 0.1499 0.17 x 10 6 19 0.3285 0.11 x 10 6 20 0.7944 0.07 x 10 6 17 0.6053 0.06 x 10 6 15

CPS/USAR TABLE 2.5-19 (Cont'd) CHAPTER 02 2.5-171 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) P-38 573.9 ML 0.0064 2.69 x 10 6 8 (Lacustrine Deposit) 0.0141 2.40 x 10 6 6 0.0269 2.15 x 10 6 9 0.0566 1.66 x 10 6 10 0.0839 1.50 x 10 6 10 0.1454 1.19 x 10 6 12 0.2789 0.90 x 10 6 16 0.7643 0.51 x 10 6 1.5905 0.29 x 10 6 10 P-38 558.9 14.1 128 0.0061 0.99 x 10 6 17 (Pre-Illinoian Glacial Till) 0.0187 0.72 x 10 6 17 0.0305 0.57 x 10 6 18 0.0575 0.44 x 10 6 17 0.1134 0.32 x 10 6 18 0.2225 0.23 x 10 6 19 0.4548 0.15 x 10 6 18 0.7560 0.12 x 10 6 18 1.5318 0.09 x 10 6 19 CPS/USAR CHAPTER 02 2.5-172 REV. 11, JANUARY 2005 TABLE 2.5-20 DYNAMIC TRIAXIAL COMPRESSION TEST DATA DAM SITE BORINGS COHESIVE SOILS BORING NUMBER ELEVATION (ft-in.) SOIL TYPE FIELD MOISTURE CONTENT (percent) DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent)

MODULUS OF RIGIDITY (lb/ft²)

DAMPING (percent) D-6 577.0 ML 8.9 135 0.0128 3.04 x 10 6 15 0.0257 2.46 x 10 6 19 (Illinoian Glacial Till) 0.0489 1.79 x 10 6 23 0.1003 1.26 x 10 6 26 0.2383 1.02 x 10 6 24 0.4767 0.46 x 10 6 26 1.2858 0.22 x 10 6 24 1.8817 0.20 x 10 6 24 3.6380 0.11 x 10 6 20 3.5125 0.13 x 10 6 19 D-11 609.8 ML 6.0 143 0.0262 6.92 x 10 6 11 0.0444 5.86 x 10 6 12 (Illinoian Glacial Till) 0.0523 4.96 x 10 6 13 0.131 3.84 x 10 6 16 0.263 2.40 x 10 6 18 0.339 2.20 x 10 6 19 0.664 1.39 x 10 6 21 1.043 1.08 x 10 6 22 1.340 0.99 x 10 6 -- D-11 599.8 CL 7.9 139 0.0097 6.01 x 10 6 34 0.0237 3.50 x 10 6 19 (Illinoian Glacial Till) 0.0473 2.47 x 10 6 21 0.0947 1.75 x 10 6 24 0.2491 0.88 x 10 6 27 0.4982 0.54 x 10 6 27 0.9964 0.33 x 10 6 24 1.8683 0.10 x 10 6 21 3.6121 0.12 x 10 6 16 CPS/USAR CHAPTER 02 2.5-173 REV. 11, JANUARY 2005 TABLE 2.5-21 DYNAMIC TRIAXIAL COMPRESSION TEST DATA DAM SITE BORINGS COHESIVE SOILS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

RELATIVE DENSITY (percent)

SINGLE\ AMPLITUDE SHEAR STRAIN (percent)

MODULUS OF RIGIDITY (lb/ft²)

DAMPING (percent) D-11 464.8 SP 19.5 108 80 0.0043 6.90 x 10 6 13 0.0101 5.42 x 10 6 10 (Mahomet Bedrock Valley Deposit) 0.0222 5.37 x 10 6 6 0.0468 4.51 x 10 6 5 0.0889 4.20 x 10 6 5 0.2402 2.38 x 10 6 12 0.4684 1.66 x 10 6 18 0.9008 1.31 x 10 6 20 1.8016 0.76 x 10 6 19 3.3630 0.13 x 10 6 27 D-11 444.8 SP 19.3 108 80 0.0451 4.01 x 10 6 11 0.0938 2.25 x 10 6 9 (Mahomet Bedrock Valley Deposit) 0.157 1.71 x 10 6 9 0.251 1.52 x 10 6 10 0.425 1.16 x 10 6 17 0.543 1.10 x 10 6 19 0.853 0.94 x 10 6 20 1.379 0.74 x 10 6 24 1.652 0.72 x 10 6 20 D-11 424.8 SP 19.4 112 82 0.0067 6.63 x 10 6 12 0.0087 6.87 x 10 6 11 (Mahomet Bedrock Valley Deposit) 0.0259 4.80 x 10 6 8 0.0506 4.17 x 10 6 7 0.0964 3.99 x 10 6 6 0.3616 1.58 x 10 6 11 0.4821 2.06 x 10 6 15 0.9643 1.22 x 10 6 18 1.9286 0.87 x 10 6 19 CPS/USAR CHAPTER 02 2.5-174 REV. 11, JANUARY 2005 TABLE 2.5-22 DYNAMIC TRIAXIAL COMPRESSION TEST DATA ULTIMATE HEAT SINK BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) H-6 619.3 ML 8.6 136 0.0054 4.03 x 10 6 8 (Illinoian Glacial Till) 0.0165 2.99 x 10 6 12 0.0346 2.26 x 10 6 16 0.0757 1.51 x 10 6 18 0.1460 1.02 x 10 6 21 0.2974 0.66 x 10 6 22 0.6083 0.41 x 10 6 20 1.0247 0.28 x 10 6 18 H-14 635.3 ML 9.5 140 0.0134 0.39 x 10 6 18 (Illinoian Glacial Till) 0.0263 0.28 x 10 6 18 0.0407 0.22 x 10 6 18 0.0945 0.13 x 10 6 16 0.1870 0.08 x 10 6 16 0.3019 0.06 x 10 6 16 0.6169 0.04 x 10 6 15 H-20 706.8 ML 8.6 136 0.0076 1.83 x 10 6 11 (Wisconsinan Glacial Till) 0.0209 1.25 x 10 6 14 0.0449 0.85 x 10 6 15 0.1015 0.53 x 10 6 15 0.1519 0.54 x 10 6 15 0.1875 0.44 x 10 6 15 0.1875 0.54 x 10 6 13 0.3332 0.37 x 10 6 14 0.4068 0.31 x 10 6 14 0.4068 0.36 x 10 6 13 0.7361 0.29 x 10 6 12 CPS/USAR TABLE 2.5-22 (Cont'd) CHAPTER 02 2.5-175 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) 0.8330 0.26 x 10 6 12 0.8330 0.29 x 10 6 12 H-20 686.8 ML 0.0065 1.65 x 10 6 -- (Interglacial Soil) 0.0157 1.24 x 10 6 13 0.0373 0.92 x 10 6 16 0.0853 0.62 x 10 6 20 0.1630 0.45 x 10 6 21 0.3324 0.30 x 10 6 21 0.7055 0.18 x 10 6 21 H-20 672.3 ML 10.5 136 0.0072 2.33 x 10 6 13 0.0165 1.84 x 10 6 14 (Illinoian Glacial Till) 0.O336 1.39 x 10 6 16 0.0804 0.82 x 10 6 19 0.1608 0.49 x 10 6 18 0.3373 0.26 x 10 6 17 0.6918 0.13 x 10 6 15 H-23 677.8 ML 9.1 138 0.0035 5.46 x 10 6 33 (Illinoian Glacial Till) 0.0096 3.24 x 10 6 31 0.0257 1.96 x 10 6 28 0.0600 1.20 x 10 6 26 0.1281 0.71 x 10 6 24 0.2835 0.41 x 10 6 22 0.5948 0.22 x 10 6 -- 0.9967 0.16 x 10 6 20 H-30 638.5 ML 6.8 145 0.0027 5.85 x 10 6 8 (Illinoian Glacial Tilll) 0.0084 4.76 x 10 6 9 0.0219 3.50 x 10 6 12 CPS/USAR TABLE 2.5-22 (Cont'd) CHAPTER 02 2.5-176 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) 0.0519 2.21 x 10 6 16 0.1082 1.34 x 10 6 -- 0.2597 0.90 x 10 6 -- 0.5518 0.60 x 10 6 -- H-36 622.7 ML 9.4 137 0.0051 4.58 x 10 6 10 (Illinoian Glacial Tilll) 0.0159 3.25 x 10 6 13 0.0325 2.55 x 10 6 16 0.0716 1.73 x 10 6 19 0.1386 1.15 x 10 6 18 0.2721 0.76 x 10 6 18 0.5881 0.56 x 10 6 -- 0.5754 0.57 x 10 6 -- 0.9819 0.50 x 10 6 -- H-9* 687.1 ML 9.0 126.2 0.0019 5.05 x 10 6 11 (Interglacial Soil) 0.0246 1.40 x 10 6 13 0.1120 0.62 x 10 6 17 0.6470 0.19 x 10 6 20 H-15* 706.3 ML 14.7 118.5 0.0157 0.76 x 10 6 4 0.0470 0.48 x 10 6 17 0.2310 0.23 x 10 6 21 1.1000 0.08 x 10 6 15 H-32* 656.1 ML 7.9 137.3 0.0076 2.14 x 10 6 5 0.0360 0.91 x 10 6 17 0.1700 0.51 x 10 6 25 0.6920 0.18 x 10 6 27 ________________________________

• Saturated, consolidated-undrained test CPS/USAR CHAPTER 02 2.5-177 REV. 11, JANUARY 2005 TABLE 2.5-23 DYNAMIC TRIAXIAL COMPRESSION TEST DATA SECTION E-E' ALONG NORTH FORK OF SALT CREEK BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) D-31 652.7 ML 7.1 140 0.0275 2.61 x 10 6 9 0.0546 1.76 x 10 6 10 (Illinoian Glacial Till) 0.0934 1.49 x 10 6 9 0.140 1.31 x 10 6 10 0.192 1.10 x 10 6 13 0.451 0.79 x 10 6 19 0.672 0.59 x 10 6 19 1.030 0.42 x 10 6 21 D-31 643.7 ML 7.6 140 0.0247 1.72 x 10 6 26 0.0494 1.13 x 10 6 29 (Illinoian Glacial Till) 0.0913 0.89 x 10 6 30 0.2536 0.42 x 10 6 29 0.5072 0.23 x 10 6 30 1.3315 0.13 x 10 6 26 2.0290 0.12 x 10 6 20 3.4239 0.11 x 10 6 6 5.0725 0.09 x 10 6 13 D-31 538.7 ML 9.2 134 0.0266 3.88 x 10 6 15 0.0516 2.64 x 10 6 13 (Illinoian Glacial Till) 0.0840 2.04 x 10 6 9 0.1284 2.00 x 10 6 9 0.1652 1.21 x 10 6 15 0.404 0.89 x 10 6 19 0.824 0.53 x 10 6 21 CPS/USAR TABLE 2.5-23 (Cont'd) CHAPTER 02 2.5-178 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE FIELD MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) 1.135 0.45 x 10 6 20 1.596 0.37 x 10 6 19 D-31 498.7 CL-ML 17.5 114 0.0121 4.01 x 10 6 9 0.0253 2.60 x 10 6 16 (Mahomet Bedrock Valley Deposit - Silty Alluvium) 0.0456 2.11 x 10 6 16 0.1215 1.28 x 10 6 14 0.2532 0.84 x 10 6 16 0.5063 0.58 x 10 6 18 0.9810 0.37 x 10 6 21 1.9620 0.19 x 10 6 25 3.6076 0.13 x 10 6 32 4.8101 0.08 x 10 6 38 CPS/USAR CHAPTER 02 2.5-179 REV. 11, JANUARY 2005 TABLE 2.5-24 DYNAMIC TRIAXIAL COMPRESSION TEST DATA STRUCTURAL FILL BORROW MATERIAL BORING NUMBER ELEVATION (ft) SOIL TYPE REMOLDED DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) CONFINING PRESSURE (lb/ft²) SINGLE AMPLITUDE SHEAR STRAIN (percent)

MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) Combined bulk sample: (Salt Creek Alluvium)

SM 123 13.5 6800 0.0074 2.186 x 10 6 8 (80% Relative 0.0187 2.200 x 10 6 9 G-18 663 to 654 Density) 0.0275 2.136 x 10 6 10 G-19 673 to 663 0.0669 1.555 x 10 6 12 G-20 657 to 647 0.0983 1.292 x 10 6 13 0.1865 0.844 x 10 6 16 0.4205 0.356 x 10 6 17 0.7758 0.185 x 10 6 16 SM 123 13.5 9700 0.0121 3.193 x 10 6 10 (80% Relative 0.0159 3.253 x 10 6 12 Density) 0.0306 2.609 x 10 6 12 0.0606 1.768 x 10 6 12 0.2073 0.535 x 10 6 16 0.4329 0.233 x 10 6 17 0.9041 0.851 x 10 6 15 SM 129 12.0 6800 0.0064 3.738 x 10 6 1 (90% Relative 0.0174 3.101 x 10 6 9 Density) 0.0331 2.496 x 10 6 12 0.0726 1.673 x 10 6 13 0.2107 0.607 x 10 6 16 0.4448 0.244 x 10 6 16 SM 129 12.0 9700 0.0074 3.894 x 10 6 (90% Relative 0.0156 3.455 x 10 6 8 Density) 0.0280 3.114 x 10 6 10 0.0684 2.031 x 10 6 11 0.1049 1.648 x 10 6 12 0.1827 1.007 x 10 6 16 0.4419 0.315 x 10 6 16 0.9045 0.130 x 10 6 12 CPS/USAR CHAPTER 02 2.5-180 REV. 11, JANUARY 2005 TABLE 2.5-25 DYNAMIC TRIAXIAL COMPRESSION TEST DATA REMOLDED WISCONSINAN TILL BORING NUMBER ELEVATION (ft) SOIL TYPE REMOLDED DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) CONFINING PRESSURE (lb/ft²)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) Combined sample from:

P-30 CL 120.3 8.7 600 0.0163 1.237 x 10 6 11 P-32 to 0.0368 1.288 x 10 6 9 P-35 727 to 692 0.0772 1.108 x 10 6 10 P-37 0.1152 1.088 x 10 6 10 P-39 0.2137 0.905 x 10 6 11 P-42 CL 118.6 8.7 1815 0.0063 3.583 x 10 6 13 0.0148 3.028 x 10 6 13 0.0315 2.363 x 10 6 15 0.0755 1.595 x 10 6 12 0.1126 1.520 x 10 6 12 0.2013 1.248 x 10 6 12 CL 121.5 8.5 3630 0.0060 3.893 x 10 6 11 0.0128 3.528 x 10 6 11 0.0281 2.725 x 10 6 12 0.0693 1.975 x 10 6 14 0.1124 1.479 x 10 6 12 0.2128 1.135 x 10 6 11 CL 124.8 8.5 635 0.0078 2.100 x 10 6 10 0.0174 1.621 x 10 6 14 0.0410 1.032 x 10 6 12 CPS/USAR TABLE 2.5-25 (Cont'd) CHAPTER 02 2.5-181 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft) SOIL TYPE REMOLDED DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) CONFINING PRESSURE (lb/ft²)

SINGLE AMPLITUDE SHEAR STRAIN (percent) MODULUS OF RIGIDITY (lb/ft²) DAMPING (percent) Combined sample from:

0.0831 9.585 x 10 6 8 P-30 0.1614 9.794 x 10 6 9 P-32 to 0.3313 9.273 x 10 6 9 P-35 727 to 692 P-37 CL 126.1 8.5 1905 0.0061 4.261 x 10 6 12 P-39 0.0164 3.273 x 10 6 11 P-42 0.1581 2.244 x 10 6 7 0.3276 2.039 x 10 6 8 0.6753 1.641 x 10 6 10 CL 126.6 8.5 3810 0.0422 4.107 x 10 6 -- 0.0809 3.345 x 10 6 8 0.1570 2.877 x 10 6 9 0.3193 2.332 x 10 6 10 0.6738 1.653 x 10 6 11 CPS/USAR CHAPTER 02 2.5-182 REV. 11, JANUARY 2005 TABLE 2.5-26 RESONANT COLUMN TEST DATA STATION SITE BORING BORING NUMBER ELEVATION (ft-in.)

SOIL OR ROCK TYPE CONFINING PRESSURE (lb/ft²) MODULUS OF RIGIDITY (lb/ft²)

DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) P-14 609.3 ML 1,000 4.83 x 10 6 139 7.6 2,000 6.16 x 10 6 (Illinoian Glacial Till) 4,300 7.64 x 10 6 8,600 9.26 x 10 6

P-14 468.3 Siltstone 0 67.81 x 10 6 144 3.0 5,000 78.76 x 10 6 (Pennsylvanian Age- 7,000 78.92 x 10 6 Modesto Formation) 10,000 78.92 x 10 6 P-18 480.2 Limestone 0 81.22 x 10 6 167 2.0 4,000 138.83 x 10 6 6,000 143.49 x 10 6 8,000 146.93 x 10 6

P-32 518.4 ML 7,200 5.51 x 10 6 120 14.4 14,400 5.99 x 10 6 (Pre-Illinoian Lacustrine 21,600 6.05 x 10 6 Deposit) 28,800 6.24 x 10 6 36,000 6.24 x 10 6

P-36 569.2 ML 7,200 3.33 x 10 6 133 9.5 (Pre-Illinoian Glacial Till) 14,400 3.80 x 10 6 21,600 3.98 x 10 6 28,800 4.06 x 10 6

CPS/USAR TABLE 2.5-26 (Cont'd) CHAPTER 02 2.5-183 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL OR ROCK TYPE CONFINING PRESSURE (lb/ft²) MODULUS OF RIGIDITY (lb/ft²)

DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) P-38 584.9 ML 4,220 8.79 x 10 6 139 8.2 (Illinoian Glacial Till) 15,850 10.59 x 10 6 24,500 12.06 x 10 6 P-38 574.9 ML 7,200 4.48 x 10 6 132 11.0 (Lacustrine Deposit) 14,400 4.83 x 10 6 21,600 5.09 x 10 6 28,800 5.14 x 10 6

P-36 499.5 Limestone 7,200 12.30 x 10 6 166 - (Pennsylvanian Age - 14,400 13.26 x 10 6 Modeston Formation) 21,600 25.46 x 10 6 28,800 26.58 x 10 6 36,000 27.17 x 10 6

CPS/USAR CHAPTER 02 2.5-184 REV. 11, JANUARY 2005 TABLE 2.5-27 RESONANT COLUMN TEST DATA ULTIMATE HEAT SINK BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL OR ROCK TYPE CONFINING PRESSURE (lb/ft²) MODULUS OF RIGIDITY (lb/ft²)

DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) H-31 667.1 ML 7,200 7.43 x 10 6 138* ---* (Illinoian Glacial Till) 14,400 11.56 x 10 6 139* 8.7* H-6 425.8 CH 21,600 3.95 x 10 6 120 15.2 (Pre-Illinoian 28,800 4.00 x 10 6 -- -- Lacustrine Deposit) 36,000 4.00 x 10 6 -- --

H-15 691.3 CL 7,200 2.35 x 10 6 102* -- (Interglacial 14,400 3.97 x 10 6 105* 17.9* Deposit)

H-17 641.6 ML 2,045 3.93 x 10 6 139 8.7 (Illinoian 4,090 4.38 x 10 6 -- -- Glacial Till) 6,134 4.84 x 10 6 -- -- 8,180 6.38 x 10 6 -- -- H-20 716.8 ML 7,200 6.78 x 10 6 132* -- (Wisconsinan 14,400 11.39 x 10 6 134* 10.3* Glacial Till) H-22 540.8 ML 7,200 5.63 x 10 6 128 12.3 (Pre-Illinoian 14,400 6.65 x 10 6 -- -- Glacial Till) 21,600 7.12 x 10 6 -- --

H-30 633.5 ML 6,150 7.07 x 10 6 141 7.4 (Illinoian 8,194 6.03 x 10 6 -- -- Glacial Till) 10,238 6.34 x 10 6 -- -- 12,269 6.91 x 10 6 -- --

H-36 597.2 ML 8,194 5.67 x 10 6 139 7.7 (Illinoian 10,238 6.05 x 10 6 -- -- Glacial Till) 12,283 6.29 x 10 6 -- --

_________________________

• Saturated, consolidated-undrained test CPS/USAR CHAPTER 02 2.5-185 REV. 11, JANUARY 2005 TABLE 2.5-28 RESONANT COLUMN TEST DATA*

STRUCTURAL FILL BORROW BORING NUMBER ELEVATION (ft) SOIL OR ROCK TYPE CONFINING PRESSURE (lb/ft²) MODULUS OF RIGIDITY (lb/ft²)

DRY DENSITY (lb/ft³) MOISTURE CONTENT (percent) Combined bulk SM 4800 29.26 x 10 5 123.1 Dry sample from: (78% 6200 41.63 x 10 5 Relative G-18 663-654 Density) G-19 673-663 7400 47.54 x 10 5 G-20 657-647 9000 53.10 x 10 5

SM 4800 35.10 x 10 5 128.9 Dry (Salt Creek (90% Alluvium) 6200 41.13 x 10 5 Relative Density) 7400 47.96 x 10 5 9000 54.65 x 10 5 9648 57.03 x 10 5

__________________________________

• Tests performed at 0.001 percent shear strain.

CPS/USAR CHAPTER 02 2.5-186 REV. 11, JANUARY 2005 TABLE 2.5-29 RESONANT COLUMN TEST DATA REMOLDED WISCONSINAN TILL*

BORING NUMBER ELEVATION SOIL OR ROCK TYPE CONFINING PRESSURE (lb/ft²) MODULUS OF RIGIDITY (lb/ft²)

DRY DENSITY (lb/ft³) COMPACTION EFFORT**

(percent) MOISTURE CONTENT (percent) Combined sample from:

P-30 CL 600 3.56 x 10 6 119.9 89.2 8.7 P-32 to 1814 5.26 x 10 6 P-35 727 to 692 3628 6.34 x 10 6 P-37 5440 7.13 x 10 6 P-39 P-42

CL 1814 5.48 x 10 6 124.3 92.5 8.5 3628 7.08 x 10 6 5440 7.96 x 10 6

____________________________

• As compacted. ** AASHO Test Designation T-180.

CPS/USAR CHAPTER 02 2.5-187 REV. 11, JANUARY 2005 TABLE 2.5-30 SHOCKSCOPE TEST DATA STATION SITE BORINGS BORING NUMBER ELEVATION (ft-in) SOIL OR ROCK TYPE GEOLOGIC UNIT CONFINING PRESSURE (lb/ft²) VELOCITY OF COMPRESSIONAL WAVE PROPAGATION (ft/sec) P-14 599.3 ML Illinoian Glacial Till 0 6666 P-14 529.3 CL Pre-Illinoian Glacial Till 0 6553 P-14 460.3 Shale Pennsylvanian Age-Modesto Formation 0 6423 P-18 704.2 ML Wisconsinan Glacial Till 0 6321 P-18 485.2 Limestone Pennsylvanian Age-Modesto Formation 0 7776 P-18 468.2 Shale Pennsylvanian Age-Modesto Formation 0 5280 CPS/USAR CHAPTER 02 2.5-188 REV. 11, JANUARY 2005 TABLE 2.5-31 LABORATORY PERMEABILITY TEST DATA STATION SITE BORINGS BORING ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT TYPE OF TEST FIELD MOISTURE CONTENT (percent)

FIELD DRY DENSITY (lb/ft³) AVERAGE COEFFICIENT OF PERMEABILITY AT 20° C K (cm/sec) P-14 654.8 ML Illinoian Glacial Till Falling Head 9.5 129 2.5 x 10

-8 P-14 579.8 ML Illinoian Glacial Till Falling Head 8.1 139 9.5 x 10

-9 P-18 683.7 ML.SM Illinoian Glacial Till Falling Head 10.3 131 2.3 x 10

-7 CPS/USAR CHAPTER 02 2.5-189 REV. 11, JANUARY 2005 TABLE 2.5-32 LABORATORY PERMEABILITY TEST DATA DAM SITE BORINGS BORING ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT TYPE OF TEST FIELD MOISTURE CONTENT (percent)

FIELD DRY DENSITY (lb/ft³) AVERAGE COEFFICIENT OF PERMEABILITY AT 20° C K (cm/sec) ESTIMATED POROSITY (percent) D-3 626.2 ML Illinoian Glacial Till Falling Head 7.5 144 3.9 x 10

-9 16.8 D-10 627.0 ML Illinoian Glacial Till Falling Head 7.2 131 1.0 x 10

-8 16.3 D-13 676.4 SP Interglacial Zone Constant Head 24.8 94 1.8 x 10

-4 40.0 SP D-13 661.4 SW Interglacial

Zone Constant

Head 6.4 105 4.7 x 10

-3 14.8 D-13 632.0 ML Illinoian Glacial Till Falling Head 7.3 142 3.8 x 10

-9 16.4 D-24 631.0 ML Salt Creek Alluvium Falling Head 7.4 123 1.8 x 10

-8 16.5 SP D-34 664.8 GP Interglacial

Zone Constant

Head 6.2 112 2.3 x 10

-3 14.3 SP D-34 649.8 GP Interglacial

Zone Constant

Head 17.5 118 2.0 x 10

-4 32.0 CPS/USAR TABLE 2.5-32 (Cont'd) CHAPTER 02 2.5-190 REV. 11, JANUARY 2005 BORING ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT TYPE OF TEST FIELD MOISTURE CONTENT (percent)

FIELD DRY DENSITY (lb/ft³) AVERAGE COEFFICIENT OF PERMEABILITY AT 20° C K (cm/sec) ESTIMATED POROSITY (percent) D-34 629.8 ML Illinoian Glacial Till Falling Head 7.8 138 6.5 x 10

-9 17.4 SP D-37 663.7 SW Interglacial

Zone Constant

Head 12.2 116 3.0 x 10

-3 24.7 ML D-37 643.7 CL Illinoian Glacial

Till Falling Head 11.7 134 1.3 x 10

-8 24.0 CPS/USAR CHAPTER 02 2.5-191 REV. 11, JANUARY 2005 TABLE 2.5-33 LABORATORY PERMEABILITY TEST DATA REMOLDED SAMPLES***

REMOLDED DATA AVERAGE COEFFICIENT BORING ELEVATION (ft-in.)

SOIL TYPE* GEOLOGIC UNIT TYPE OF TEST MOISTURE CONTENT (percent)

DRY DENSITY (lb/ft³) COMPACTION EFFORT**

(percent)

OF PERMEABILITY AT 20° C K (cm/sec) 702.6 to S-10 697.6 CL Wisconsinan Glacial Till Falling Head 13.6 126 97.3 8.2 x 10

-9 702.6 to S-10 697.6 CL Wisconsinan Glacial Till Falling Head 12.4 125 96.5 2.0 x 10

-8 727.2 to S-14 720.2 CL Wisconsinan Glacial Till Falling Head 16.8 109 91.5 1.6 x 10

-8 727.2 to S-14 720.2 CL Wisconsinan Glacial Till Falling Head 11.0 125 105.0 1.0 x 10

-8

___________________________

• See Figure 2.5-355 for definition of soil type symbols.

• • AASHO Test Designation T-180.

• • • As compacted.

CPS/USAR CHAPTER 02 2.5-192 REV. 11, JANUARY 2005 TABLE 2.5-34 RELATIVE DENSITY TEST DATA BORING NUMBER ELEVATION (ft-in.) GEOLOGIC UNIT MINIMUM DRY DENSITY (lb/ft³) MAXIMUM DRY DENSITY (lb/ft³)

IN SITU RELATIVE DENSITY (percent) D-11 473.8 Mahomet Bedrock Valley Deposit 92 113 (Wet Method)-- D-11 424.8 Mahomet Bedrock Valley Deposit 91 118 (Wet Method)82 CPS/USAR CHAPTER 02 2.5-193 REV. 11, JANUARY 2005 TABLE 2.5-35 RESULTS OF STRESS-CONTROLLED CYCLIC TRIAXIAL (LIQUEFACTION) TESTS NUMBER OF CYCLES REQUIRED TO CAUSE SAMPLE NUMBER SOIL DESCRIPTION*

DRY DENSITY MOLDING MOISTURE CONTENT (percent) PRINCIPAL CONSOLIDATION STRESS RATIO, Kc EFFECTIVE CONFINING PRESSURE (psf) CYCLIC STRESS RATIO cv2/ SKEMPTON'S PORE PRESSURE PARAMETER B INITIAL LIQUE- FACTION***

5% STRAIN** Type B Granular Structural Fill Material S-1 Light brown fine 126.5 7.0 1.0 6800 0.403 0.95 33 42 S-2 to coarse sand 126.4 7.0 1.0 6800 0.504 0.95 20 33 S-3 with gravel 127.2 7.0 1.0 6800 0.655 0.96 13 12

S-4 127.4 7.0 1.0 6800 0.302 0.95 50 55

__________________________

• For soil properties see Figure 2.5-348.

• • Double Amplitude Axial Strain. *** Initial liquefaction is defined as when the increase in pore pressure is equal to the effective confining pressure.

CPS/USAR CHAPTER 02 2.5-194 REV. 11, JANUARY 2005 TABLE 2.5-36 AVERAGE SOIL PROPERTIES OF BORROW MATERIALS FOR THE ULTIMATE HEAT SINK

SOIL PROPERTIES

Soil Type Brown and gray clayey silt to silty clay with trace to some sand and trace fine gravel.

AVERAGE Liquid Limit (%) 21.7 Plasticity Index (%) 7.7 Compaction Test (ASTM D1557) Maximum Dry Density (lb/ft

3) 133.8 Optimum Moisture Content (%) 8.4 CPS/USAR CHAPTER 02 2.5-195 REV. 11, JANUARY 2005 TABLE 2.5-37 AVERAGE SOIL PROPERTIES OF COHESIVE BORROW MATERIALS FOR THE MAIN PLANT AND SCREEN HOUSE SOIL PROPERTIES

Soil Type Brown and gray silty clay with trace to some sand and trace fine gravel.

AVERAGE Liquid Limit (%) 23.4 Plasticity Index (%) 8.9 Compaction Test (ASTM D1557) Maximum Dry Density (lb/ft

3) 133.1 Optimum Moisture Content (%) 8.7 CPS/USAR CHAPTER 02 2.5-196 REV. 11, JANUARY 2005 TABLE 2.5-38 FIELD PERMEABILITY TESTS BORING GROUND SURFACE ELEVATION (ft-in.)

ZONE OF PERCOLATION ELEVATION (ft-in.) GEOLOGIC UNIT AVERAGE COEFFICIENT OF PERMEABILITY, k (cm/sec) ESTIMATED POROSITY (percent) D-19 658.9 625.0 to 620.9 Illinoian Till 1.4 x 10-5 26.7 D-23 655.8 630.8 to 624.3 Illinoian Till 6.1 x 10

-6 24.5 E-1B 733.0 703.0 to 693.0 Wisconsinan Till 1.5 x 10 -

P-37 741.5 726.1 to 701.1 Wisconsinan Till 2.6 x 10

-6 25.7 CPS/USAR CHAPTER 02 2.5-197 REV. 11, JANUARY 2005 TABLE 2.5-39 OBSERVED SURFACE WAVE CHARACTERISTICS STATION SITE OBSERVED WAVE WAVE TYPE PREDOMINANT PARTICLE MOTION PREDOMINANT FREQUENCY (Hz) APPARENT WAVELENGTH (ft) APPARENT VELOCITY (ft/sec) OBSERVED LENGTH OF WAVE TRAIN (cycles) 1 Rayleigh Transverse -

Radial 11 - 12 240 2900 6 2 Sezawa Transverse -

M2 Vertical 6 - 7 240 1500 - 1600 5 3 Sezawa Transverse -

M2 Radial 7 120 - 140 900 8 - 10 4 (*Rayleigh?) Vertical - greater Radial 7 80 - 100 than -- 600 - 700

_______________________

• There is an indication of a Rayleigh wave having a velocity greater than 600 - 700 feet per second and a frequency of 7 Hz.

The initial motion of this wave could not be determined from the field data.

CPS/USAR CHAPTER 02 2.5-198 REV. 11, JANUARY 2005 TABLE 2.5-40 OBSERVED SURFACE WAVE CHARACTERISTICS DAM SITE OBSERVED WAVE WAVE TYPE PREDOMINANT PARTICLE MOTION PREDOMINANT FREQUENCY (Hz) APPARENT WAVE LENGTH (ft) APPARENT VELOCITY (ft/sec) OBSERVED LENGTH OF WAVE TRAIN (cycles) 1 Sezawa Radial -

M2 Vertical 10 190 1900 7 2 Rayleigh Radial - Transverse 8 125 1000 6 3 ?* Radial - Transverse 7 - 8 50 - 60 400 - 500 8

___________________________

• This wave displays little motion on the vertical component; however, this could be the result of normal attenuation, and not the indication of a Love wave.

CPS/USAR CHAPTER 02 2.5-199 REV. 11, JANUARY 2005 TABLE 2.5-41 OBSERVED SURFACE WAVE CHARACTERISTICS SECTION E - E' ALONG NORTH FORK OF SALT CREEK OBSERVED WAVE WAVE TYPE PREDOMINANT PARTICLE MOTION PREDOMINANT FREQUENCY (Hz) APPARENT WAVE LENGTH (ft) APPARENT VELOCITY (ft/sec) OBSERVED LENGTH OF WAVE TRAIN (cycles) 1 Sezawa Verticle -

10 150 1500 8 - 10 M2 Transverse 2 Sezawa Radial - 10 70 700 6 M2 Transverse CPS/USAR CHAPTER 02 2.5-200 REV. 11, JANUARY 2005 TABLE 2.5-42 AMBIENT GROUND MOTION MEASUREMENTS (June 6 and 13, 1972) GROUND MOTION** LOCATION TIME FREQUENCY (Hz) GAIN* (x 100) RAD VERT TRANS*** Station site:

V x 5 0.06 0.06 0.06 Boring 9:30 AM 10-12, 50,150 A x 20 0.12 0.08 0.10 P-18 June 6 D x 20 0.0015 0.002 0.002 Station site:

V x 20 0.05 0.04 0.08 Boring 5:00 PM 10-12, 50 A x 20 0.058 0.058 0.067 P-28 June 6 D x 20 0.0015 0.0005 0.0010 Dam site: Boring 10:30 AM 8-10,33-1/3, V x 20 0.07 0.10 0.11 D-11 June 6 50, 100 A x 20 0.067 0.083 0.12 D x 20 0.003 0.0005 0.001 Section E - E' along the North Fork of Salt Creek: Boring 2:00 PM 4, 25,33-1/3, V x 20 0.08 0.08 0.09 D-31 June 13 150 A x 20 0.017 0.05 0.033 D x 20 0.001 0.001 0.001 ________________________________

• V = velocity (in./sec)

A = acceleration (in./sec/sec)

D = displacement (in.).

• • All values are x 10

-3. *** RAD = radial VERT = vertical TRANS = transverse.

CPS/USAR CHAPTER 02 2.5-201 REV. 11, JANUARY 2005 TABLE 2.5-43 LABORATORY COMPRESSIONAL WAVE VELOCITY TABULATION BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN %VELOCITY (ft/sec) P-14 713.8 ML Wisconsinan Dynamic 2,500 0.0141320 (24.5) Glacial Triaxial 0.02641220 Till Compression 0.05211090 Test 0.1268880 0.2508740 0.5191690 1.2540390 1.9153358 2.5205337

P-15 691.8 ML Interglacial Dynamic 3,000 0.01441280 (44.5) Zone Triaxial 0.02771165 Compression 0.05311085 Test 0.1037920 0.2529740 0.5059605 1.2647456 2.5170354 3.6941290 5.0841260

P-14 624.3 CL Illinoian Dynamic 9,000 0.01271830 (114.0) Glacial Triaxial 0.02621610 Till Compression 0.05081375 Test 0.12721085 0.2502930 0.5077775 0.9921675 1.9841550 3.8697483 CPS/USAR TABLE 2.5-43 (Cont'd) CHAPTER 02 2.5-202 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN %VELOCITY (ft/sec) P-22 659.5 ML Illinoian Dynamic 6,000 0.01251840 (74.5) Glacial Triaxial 0.02511560 Till Compression 0.04941380 Test 0.12451130 0.2462950 0.4803700 0.7217676 0.9721620

D-6 577.00 ML Illinoian Dynamic 6,000 0.01281975 (79.0) Glacial Triaxial 0.02571790 Till Compression 0.04891530 Test 0.10031275 0.23831115 0.4767770 1.2858535 1.8817510 3.5125412 3.3680378 D-11 599.8 CL Illinoian Dynamic 4,000 0.00972790 (54.0) Glacial Triaxial 0.02372140 Till Compression 0.04731790 Test 0.09471510 0.24911070 0.4982840 0.9964650 1.8683360 3.6121395

CPS/USAR TABLE 2.5-43 (Cont'd) CHAPTER 02 2.5-203 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN %VELOCITY (ft/sec) D-11 609.8 ML Illinoian Dynamic 4,000 0.02623000 (44.0) Glacial Triaxial 0.04442760 Till Compression 0.05232540 Test 0.1312230 0.2631760 0.3391690 0.6641340

1.0 431180

1.3401130

D-11 464.8 SP Mahomet Dynamic 9,000 0.00433200 (189.0) Bedrock Triaxial 0.01012840 Valley Compression 0.02222820 Deposit Test 0.04682590

0.0 8892500

0.24021880 0.46841530 0.90081370 1.80161060 3.363440

D-11 444.8 SP Mahomet Dynamic 9,000 0.04512440 (209.0) Bedrock Triaxial 0.09381830 Valley Compression 0.1591590 Deposit Test 0.2511510 0.4251310 0.5431275 0.8531184 1.3791050 1.6521030 CPS/USAR TABLE 2.5-43 (Cont'd) CHAPTER 02 2.5-204 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN %VELOCITY (ft/sec) D-31 498.7 CL Mahomet Dynamic 10,000 0.01212370 (169.0) Bedrock Triaxial 0.02531920 Valley Compression 0.04561730 Deposit Test 0.12151350 0.25321090 0.5063900 0.9810720 1.962520 3.6076425 4.8101316

D-11 424.8 SP Mahomet Dynamic 11,000 0.00672930 (229.0) Bedrock Triaxial 0.00872990 Valley Compression 0.02592500 Deposit Test 0.05062330

0.0 9642280

0.36161450 0.48211630 0.96431260 1.92861060

D-31 538.7 ML Illinoian Dynamic 10,000 0.02662240 (129.0) Glacial Triaxial 0.05161850 Till Compression 0.08401640 Test 0.12841610 0.16521255 0.4041072 0.824825 1.135761 1.596690 CPS/USAR TABLE 2.5-43 (Cont'd) CHAPTER 02 2.5-205 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN %VELOCITY (ft/sec) D-31 643.7 ML Illinoian Dynamic 2,500 0.02471480 (24.0) Glacial Triaxial 0.04941190 Till Compression 0.09131060 Test 0.2536727 0.5072537 1.3315407 2.0290398 3.4239360

5.0 725338

D-31 652.7 ML Illinoian Dynamic 2,500 0.02751840 (15.0) Glacial Triaxial 0.05461520 Till Compression 0.09341390 Test 0.1401300 0.1921200 0.4511020 0.672870 1.030736 P-14 599.3 ML Illinoian Shockscope 0 6666 (139.0) Glacial Test Till

P-14 529.3 CL Pre-Illinoian Shockscope 0 6553 (209.0) Glacial Test Till

P-14 460.3 Shale Shockscope 0 6423 (278.0) Test

CPS/USAR TABLE 2.5-43 (Cont'd) CHAPTER 02 2.5-206 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN %VELOCITY (ft/sec) P-18 704.2 ML Wisconsinan Shockscope 0 6321 (34.0) Glacial Test Till P-18 485.2 Lime- Shockscope 0 7776 (253.0) stone Test P-18 468.2 Shale Shockscope 0 5280 (270.0) Test

CPS/USAR CHAPTER 02 2.5-207 REV. 11, JANUARY 2005 TABLE 2.5-44 LABORATORY SHEAR WAVE VELOCITY TABULATION BORING NUMBER ELEVATION (depth) SOIL OR ROCK TYPE GEOLOGIC UNIT DATA SOURCE CONFINING PRESSURE (lb/ft²) SHEAR STRAIN % VELOCITY (ft/sec) P-14 468.3 Sandstone Resonant 10,000 0.077 4140 (270.0) Column 7,000 0.077 4140 5,000 0.077 4150 1.4 0.089 3840 P-18 480.2 Sandstone Resonant 8,000 0.044 5270 (258.0) Column 6,000 0.047 5210 4,000 0.046 5130 1.4 0.079 3920 P-14 609.3 ML Illinoian Resonant 8,640 0.672 1410 (129) Glacial Column 4,320 0.0857 1280 Till 2,016 0.1048 1150 1,008 0.1234 1040

CPS/USAR CHAPTER 02 2.5-208 REV. 11, JANUARY 2005 TABLE 2.5-45 FIELD COMPRESSIONAL WAVE VELOCITY TABULATION VELOCITY (ft/sec)

SOURCE MATERIAL TYPE DEPTH (feet) 2,000 Seismic Line 1 Low-velocity surface layer 0-14 5,750 Seismic Line 1 Wisconinan Till 14-56 7,500 Seismic Line 1 Illinoian Glacial Till 56-233 10,000 Seismic Line 1 Top of bedrock 233+ 2,000 Seismic Line 2 Low-velocity surface layer 0-15 5,700 Seismic Line 2 Wisconinan Till 15-46 7,500 Seismic Line 2 Illinoian Glacial Till 46-238 10,000 Seismic Line 2 Top of bedrock 238+ 2,000 Seismic Line 3 Low-velocity surface layer 0-15 5,700 Seismic Line 3 Wisconinan Till 15-55 7,500 Seismic Line 3 Illinoian Glacial Till 55-243 9,750 Seismic Line 3 Top of bedrock 243+ 2,000 Seismic Line 4 Low-velocity surface layer 0-16 5,500 Seismic Line 4 Wisconinan Till 16-55 7,250 Seismic Line 4 Illinoian Glacial Till 55-243 10,250 Seismic Line 4 Top of bedrock 243+ 2,000 Seismic Line 5 Low-velocity surface layer 0-12 5,750 Seismic Line 5 Wisconinan Till 12-48 7,500 Seismic Line 5 Illinoian Glacial Till 48-240 10,500 Seismic Line 5 Top of bedrock 240+ 2,000 Seismic Line 6 Low-velocity surface layer 0-18 7,300 Seismic Line 6 Wisconinan Till 18-150 5,800 Seismic Line 6 Illinoian Glacial Till 150-290 10,600 Seismic Line 6 Top of bedrock 290+

CPS/USAR TABLE 2.5-45 (Cont'd) CHAPTER 02 2.5-209 REV. 11, JANUARY 2005 VELOCITY (ft/sec)

SOURCE MATERIAL TYPE DEPTH (feet) 2,000 Seismic Line 7 Low-velocity surface layer 0-14 7,500 Seismic Line 7 Illinoian Glacial Till 14-195 6,000 Seismic Line 7 Bedrock Valley Outwash Deposit 195-305 9,800 Seismic Line 7 Top of bedrock 305+ 2,875 Uphole P-14 Low-velocity surface layer 0-10 4,800 Uphole P-14 Wisconinan Till 10-57 7,400 Uphole P-14 Illinoian Glacial Till 57-237 12,000 Uphole P-14 Top of bedrock 237+ 6,800 Uphole D-11 Illinoian Glacial Till 0-275 7,000 Uphole D-31 Illinoian Glacial Till 0-275 CPS/USAR CHAPTER 02 2.5-210 REV. 11, JANUARY 2005 TABLE 2.5-46 FIELD SHEAR WAVE VELOCITY TABULATION ESTIMATED VELOCITY (ft/sec)

SOURCE MATERIAL TYPE DEPTH (feet) 900* Geophysical P-14 Low-velocity surface layer 0-16 1100 Geophysical P-14 Wisconsinan till 16-47 2100 Geophysical P-14 Illinoian glacial till 47-237 5700 Geophysical P-14 Top of bedrock 237+ 900* Geophysical D-11 Salt Creek alluvium 0-18 2000-2100 Geophysical D-11 Illinoian glacial till 18-150 1800 Geophysical D-11 Bedrock valley outwash deposit 150-290 5700 Geophysical D-11 Top of bedrock 290+ 900* Geophysical D-31 Salt Creek alluvium 0-14 2100 Geophysical D-31 Illinoian glacial till 14-195 1800 Geophysical D-31 Bedrock valley 195-305 5300-5500 Geophysical D-31 Top of bedrock 305+ _______________________

• Measured.

CPS/USAR CHAPTER 02 2.5-211 REV. 11, JANUARY 2005 TABLE 2.5-47 DETAILS OF STRUCTURE FOUNDATIONS AND BEARING STRATA FOUNDATION MAT BEARING STRATUM APPROXIMATE APPROXIMATE COMPACTED SIZE, L x B BOTTOM THICKNESS STRUCTURE (ft x ft) ELEVATION (ft)TYPE* (ft)

SOIL UNDERLYING BEARING STRATUM Containment 130 dia. 702.0 B 22.0 A Fuel Building 182 x 151 702.3 B 22.0 A Auxiliary Bldg.

178 x 122 697.8 B 17.5 A Radwaste, machine Shop, and Off-Gas Bldg. 232 x 321 693.0 B 12.0 A Service Bldg.

96 x 195 732.0 B 30.0 A Diesel Generator and HVAC 222 x 106 703.0 B 22.0 A Control 219 x 100 693.0 B 12.0 A Turbine and Heater Bay 185 x 315 702.0 B 22.0 A Circulating Water Screen House 238 x 176 653.0 A - - Ultimate Heat Sink

Outlet Structure 43 x 34 669.0 C 7.0 B,A _________________________________

• A: Existing Illinoian glacial till. B: Type B structural fill (onsite granular-type material). Minimum placement relative density = 85% as determined by ASTM D2049-69 test method. C: Fly ash mixture.

CPS/USAR CHAPTER 02 2.5-212 REV. 11, JANUARY 2005 TABLE 2.5-48 PARAMETERS FOR ANALYSIS OF ROCK-SOIL-STRUCTURE INTERACTION COHESIONLESS SOIL COHESIVE SOILS RECOMPACTED RECOMPACTED WISCONSINAN WISCONSINAN GLACIAL TILL OF GLACIAL TILL OF WEDRON FORMATION WEDRON FORMATION WISCONSINAN COMPACTED TYPE A MATERIAL TYPE A MATERIAL GLACIAL TILL OF INTERGLACIAL STRUCTURAL FILL (AS COMPACTED) (SATURATED) LOESS WEDRON FORMATION DEPOSITS DENSITY (pcf): Dry density 123 127 128 101 118 115 Wet density 132 141 144 120 137 131 POISSON'S RATIO: Dynamic 0.40 0.40 0.40 0.37 0.48 0.48 Static 0.30 0.40 0.40 0.40 0.40 0.40 STATIC MODULUS OF ELASTICITY (Es)

In-situ modulus (psf)

-- 8.0 x 105 2.0 x 105 2.0 x 105 13.1 x 10 5 15.1 x 105 Increase with surcharge dEs/d'm (psf/psf) 350 0 0 0 0 0 DYNAMIC MODULUS OF ELASTICITY (psf) Single amplitude Shear strain = 1.0% 22,000 ('m)1/2 11 x 105 3 x 105 3 x 105 12 x 105 9 x 105 = 0.1% 90,000 ('m)1/2 39 x 105 8 x 105 8 x 105 36 x 105 33 x 105 = 0.01% 207,000 ('m)1/2 98 x 105 34 x 105 33 x 105 80 x 105 80 x 105 = 0.001% 271,000 ('m)1/2 148 x 105 76 x 105 74 x 105 130 x 105 130 x 105 = 0.0001% 280,000 ('m)1/2 162 x 105 95 x 105 93 x 105 160 x 105 160 x 105 STATIC MODULUS OF RIGIDITY (Gs) In-situ modulus (psf) -- 3.0 x 105 0.7 x 105 0.7 x 105 4.7 x 105 5.4 x 105 Increase with surcharge

dGs/d'm (psf/psf) 135 0 0 0 0 0 CPS/USAR TABLE 2.5-48 (Continued) CHAPTER 02 2.5-213 REV. 11, JANUARY 2005 COHESIONLESS SOIL COHESIVE SOILS RECOMPACTED RECOMPACTED WISCONSINAN WISCONSINAN GLACIAL TILL OF GLACIAL TILL OF WEDRON FORMATION WEDRON FORMATION WISCONSINAN COMPACTED TYPE A MATERIAL TYPE A MATERIAL GLACIAL TILL OF INTERGLACIAL STRUCTURAL FILL (AS COMPACTED) (SATURATED) LOESS WEDRON FORMATION DEPOSITS DYNAMIC MODULUS OF RIGIDITY (psf)

Single amplitude Shear strain = 1.0% 8,000 ('m)1/2 4 x 105 1 x 105 1 x 105 4 x 105 3 x 105 = 0.1% 32,000 ('m)1/2 14 x 105 3 x 105 3 x 105 12 x 105 11 x 105 = 0.01% 74,000 ('m)1/2 35 x 105 12 x 105 12 x 105 27 x 105 27 x 105 = 0.001% 97,000 ('m)1/2 53 x 105 27 x 105 27 x 105 44 x 105 44 x 105 = 0.0001% 100,000 ('m)1/2 58 x 105 34 x 105 34 x 105 54 x 105 54 x 105 DAMPING Percent of critical damping Single amplitude Shear strain = 1.0% 16 20 20 20 20 20 = 0.1% 14 9 15 15 9 9 = 0.01% 6 5 10 10 5 5 = 0.001% 2 3 6 6 3 3 = 0.0001% 1 2.5 4 4 2.5 2.5 CPS/USAR TABLE 2.5-48 (Continued) CHAPTER 02 2.5-214 REV. 11, JANUARY 2005 COHESIONLESS COHESIONLESS SOIL COHESIVE SOILS SOIL SALT CREEK INTERGLACIAL ILLINOIAN LACUSTRINE PRE-ILLINOIAN PRE-ILLINOIAN ALLUVIUM SAND DEPOSITS GLACIAL TILL DEPOSITS DEPOSITS DEPOSITS ROCK* DENSITY (pcf): Dry density 100 108 138 123 130 107 156 Wet density 125 120 150 134 145 126 159 POISSON'S RATIO: Dynamic 0.37 0.40 0.46 0.47 0.47 0.40 0.29 Static 0.40 0.40 0.35 0.35 0.35 0.40 0.29 STATIC MODULUS OF ELASTICITY (Es)

In-situ modulus (psf)

-- -- 43.6 x 10 5 24.9 x 10 5 42.4 x 10 5 110 x 10 5 0.7 to 3.8 x 10 8 Increase with surcharge dEs/d'm (psf/psf) 150 260 0 0 0 1100 0 DYNAMIC MODULUS OF ELASTICITY (psf) Single amplitude Shear strain = 1.0% 2,700 ('m)1/2 4,200 ('m)1/2 23 x 105 24 x 105 24 x 105 28,200 ('m)1/2 3.6 to 7.8 x 10 8 = 0.1% 11,000 ('m)1/2 17,000 ('m)1/2 88 x 105 76 x 105 76 x 105 95,000 ('m)1/2 0 = 0.01% 44,000 ('m)1/2 62,000 ('m)1/2 292 x 105 226 x 10 5 226 x 10 5 174,000 ('m)1/2 = 0.001% 52,000 ('m)1/2 81,000 ('m)1/2 496 x 105 338 x 10 5 338 x 10 5 218,000 ('m)1/2 = 0.0001% 280,000 ('m)1/2 84,000 ('m)1/2 584 x 105 412 x 10 5 412 x 10 5 238,000 ('m)1/2 STATIC MODULUS OF RIGIDITY (Gs) In-situ modulus (psf) -- -- 16.1 x 10 5 9.2 x 10 5 15.7 x 10 5 40 x 105 0.3 to 1.5 x 10 8 Increase with surcharge

dGs/d'm (psf/psf) 54 93 0 0 0 392 0 CPS/USAR TABLE 2.5-48 (Continued) CHAPTER 02 2.5-215 REV. 11, JANUARY 2005 COHESIONLESS COHESIONLESS SOIL COHESIVE SOILS SOIL SALT CREEK INTERGLACIAL ILLINOIAN LACUSTRINE PRE-ILLINOIAN PRE-ILLINOIAN ALLUVIUM SAND DEPOSITS GLACIAL TILL DEPOSITS DEPOSITS DEPOSITS ROCK* DYNAMIC MODULUS OF RIGIDITY (psf)

Single amplitude Shear strain = 1.0% 1,000 ('m)1/2 1,500 ('m)1/2 8 x 105 8 x 105 4 x 105 10,500 ('m)1/2 1.4 to 3.0 x 10 8 = 0.1% 4,000 ('m)1/2 6,000 ('m)1/2 30 x 105 26 x 105 26 x 105 34,000 ('m)1/2 0 = 0.01% 16,000 ('m)1/2 22,000 ('m)1/2 100 x 105 77 x 105 77 x 105 62,000 ('m)1/2 = 0.001% 19,000 ('m)1/2 29,000 ('m)1/2 170 x 105 115 x 10 5 115 x 10 5 78,000 ('m)1/2 = 0.0001% 20,000 ('m)1/2 30,000 ('m)1/2 200 x 105 140 x 10 5 140 x 10 5 85,000 ('m)1/2 DAMPING Percent of critical damping Single amplitude Shear strain = 1.0% 21 28 22 20 20 20 1 to 2 = 0.1% 10 13 16 9 12 10 = 0.01% 3 4 7.5 4.5 7.5 3 = 0.001% 1 1.5 4. 3 4. 2 = 0.0001% 0.5 0.5 3 2.5 3 1

• These values are valid for strain levels on the order of 10

-4 to 10-5 percent. Notes: 1. 'm = mean effective stress (psf). 2. The static modulus of elasticity values for cohesive soils were calculated based on the constrained modulus derived from th e reloading portion of the consolidation curve. 3. Pre-Illinoian cohesive deposits include glacial and lacustrine deposits. 4. Pre-Illinoian cohesionless deposits include Mahomet Valley deposits.

5. The selected parameters reflect both the results of geophysical and laboratory tests performed during this investigation an d results published and previously developed for similar soils.

CPS/USAR CHAPTER 02 2.5-216 REV. 11, JANUARY 2005 TABLE 2.5-49 SUMMARY OF LIQUEFACTION ANALYSES ELEVATION*

(FEET) EFFECTIVE OVERBURDEN PRESSURE

()2ft/kipso AVERAGE CYCLIC SHEAR STRESS FOR 10 CYCLES

()2ft/kipsliq AVERAGE CYCLIC SHEAR STRESS CAUSING LIQUE-FACTION IN 10 CYCLES

()2ft/kipsliq FACTOR OF SAFETY WITH RESPECT TO INITIAL LIQUEFACTION av/liq Elevation 730 0.79 0.13 0.42 3.18 Elevation 705 2.53 0.66 1.35 2.03 Elevation 680 4.26 1.07 2.27 2.11 _______________________________

• Grade elevation 736 feet.

CPS/USAR CHAPTER 02 2.5-217 REV. 11, JANUARY 2005 TABLE 2.5-50 SUMMARY OF LIQUEFACTION ANALYSES FOR NEW MADRID TYPE EARTHQUAKE ELEVATION*

(FEET) EFFECTIVE OVERBURDEN PRESSURE

()2ft/kipso AVERAGE CYCLIC SHEAR STRESS FOR 30 CYCLES

()2ft/kipsav AVERAGE CYCLIC SHEAR STRESS CAUSING LIQUE-FACTION IN 30 CYCLES

()2ft/kipsliq FACTOR OF SAFETY WITH RESPECT TO INITIAL LIQUEFACTION av/liq Elevation 730 .79

.07 .22 3.35 Elevation 705 2.53

.33 .71 2.14 Elevation 680 4.26

.54 1.19 2.22 _______________________________

• Grade elevation 736 feet.

CPS/USAR CHAPTER 02 2.5-218 REV. 11, JANUARY 2005 TABLE 2.5-51 HAS BEEN INTENTIONALLY DELETED.

CPS/USAR CHAPTER 02 2.5-219 REV. 11, JANUARY 2005 TABLE 2.5-52 AVERAGE SOIL PROPERTIES OF BORROW MATERIALS FOR MAIN DAM SOIL PROPERTIES Soil Type Brown and gray silty clay with trace to some sand and trace fine gravel.

AVERAGE Natural Moisture Content (%) 12.2 Liquid Limit (%) 23.6 Plasticity Index (%) 9.5 Compaction Test (ASTM D698) Maximum Dry Density (lb/ft

3) 124.1 Optimum Moisture Content (%) 11.7 Shear Strength (Remolded Samples):

Total:* Cohesion (lb/ft

2) 1300 Angle of Internal Friction (degrees) 0 Effectiveness: Cohesion (lb/ft

2) 200 Angle of Internal Friction (degrees) 33 Permeability* (cm/sec) 2x10

-8

____________________________

• Shear Strength and Permeability correspond to samples compacted to a dry density of 90% of the maximum density determined by the AASHO T-180 Method of Compaction, latest revision.

CPS/USAR CHAPTER 02 2.5-220 REV. 11, JANUARY 2005 TABLE 2.5-53 SOIL PARAMETERS FOR STATIC ANALYSIS OF NATURAL SLOPE LOESS WISCONINAN GLACIAL TILL OF WEDRON FORMATION INTER-GLACIAL DEPOSITS INTER-GLACIAL SAND DEPOSITS ILLINOIAN GLACIAL TILL OF ALTERED GLASFORD FORMATION LACUSTRINE DEPOSITS PRE-ILLINOIAN DEPOSITS Density (lb/ft

3) 120.0 137.0 131.0 125.0 150.0 134.0 145.0 Coefficient of Earth Pressure at Rest 0.5 1.2 1.2 1.0 1.0 0.6 0.7 Poisson's

Ratio 0.35 0.40 0.40 0.40 0.40 0.40 0.40 Modulus of

Elasticity (lb/ft2) 2.0 x 105 2.3 x 10 5 2.3 x 10 5 2.0 x 10 5 4.0 x 10 5 4.5 x 10 5 4.5 x 10 5 Cohesion (lb/ft2) 0 600 600 0 0 0 1400 Angle of

Internal Friction (deg.) 20 30 30 38 47 34 42 CPS/USAR CHAPTER 02 2.5-221 REV. 11, JANUARY 2005 TABLE 2.5-54 SUMMARY OF LIQUEFACTION ANALYSES BORING NUMBER ELEVATION (feet) DESCRIPTION OF SAND N-VALUE (blows/ft)

RELATIVE DENSITY(1)

(%) tavg(2) (lb/ft²)

tliq(3) (lb/ft²)

FACTOR OF SAFETY H-28 671 Gray silty fine to coarse sand, some gravel (SM) 15 80 266 200 0.75 P-5 663 Gray fine to coarse sand, some gravel (SP) 67 90 338 402 1.19 P-8 665 Gray medium to coarse sand with gravel (SP) 25 95 225 246 1.09 P-12 662 Gray fine to coarse sand, some gravel (SP) 23 91 282 286 1.01 H-32 673 Gray fine to coarse sand, some gravel, trace silt (SP) 10 68 210 121 0.57 H-33 678 Brown silty fine to coarse sand, trace clay and gravel (SW) 5 52 157 69 0.44 H-33 676 Brown fine silty sand (SM) 3 48 196 80 0.41 H-33 667 Dark gray fine sand, some silt (SP) 12 66 371 235 0.63 ______________________

(1) Relative density is obtained from the N-value and effective overburden pressure using relationship given by Gibbs and Holtz.

(2) Average cyclic shear stress for 10 cycles.

(3) Average cyclic shear stress causing liquefaction in 10 cycles. The mean grain size D 50 of all sands was assumed equal to 0.2 mm, based on available grain size analyses and soil description.

CPS/USAR CHAPTER 02 2.5-222 REV. 11, JANUARY 2005 TABLE 2.5-55 SOIL PARAMETERS FOR STATIC ANALYSIS OF SUBMERGED DIKE RECOMPACTED WISCONSINAN GLACIAL TILL OF WEDRON FORMATION TYPE A MATERIAL ILLINOIAN GLACIAL TILL OF ALTERED GLASFORD FORMATION LACUSTRINE DEPOSITS PRE-ILLINOIAN DEPOSITS Density (lb/ft

3) 135.0 150.0 134.0 145.0 Poisson's Ratio 0.40 0.35 0.35 0.35 Modulus of Elasticity (lb/ft2) 2.0 x 105 4.0 x 105 4.5 x 105 4.5 x 105 Cohesion (lb/ft

2) 400 0 0 1400 Angle of Internal Friction (deg.) 29 47 34 42 CPS/USAR CHAPTER 02 2.5-223 REV. 11, JANUARY 2005 TABLE 2.5-56 SUMMARY OF LOCAL FACTORS OF SAFETY IN SUBMERGED DIKE ELEMENTS ELEMENT NUMBER CRITICAL NUMBER OF CYCLES Nc AVERAGE CYCLIC SHEAR STRESS FOR N c CYCLES avg (lb/ft²) AVERAGE CYCLIC SHEAR STRESS CAUSING 5% STRAIN IN Nc CYCLES fail (lb/ft²) MINIMUM FACTOR OF SAFETY WITH RESPECT TO 5% STRAIN FOR N c CYCLES fail/avg 1 5 53 270 5.09 2 5 62 270 4.35 4 5 81 270 3.33 6 5 92 270 2.93 11 6 89 260 2.92 12 6 106 370 3.49 14 6 147 370 2.52 15 6 161 370 2.30 16 6 168 370 2.20 21 6 92 260 2.83 22 6 147 370 2.52 23 4 209 570 2.73 24 5 234 530 2.26 25 6 243 500 2.06 26 6 253 500 1.98 33 4 269 570 2.12 35 6 325 680 2.09 36 6 334 680 2.04 43 8 311 470 1.51 45 7 245 490 2.41 47 6 371 760 2.05 49 6 418 940 2.25 50 6 432 940 2.18 61 6 512 760 1.48 64 6 574 940 1.64 66 6 589 940 1.60 68 3 686 1160 1.69 81 6 798 940 1.18 85 6 859 940 1.09 87 2 1189 1440 1.21 89 3 963 1320 1.37 CPS/USAR CHAPTER 02 2.5-224 REV. 11, JANUARY 2005 TABLE 2.5-57 SUMMARY OF LOCAL FACTORS OF SAFETY IN THE NATURAL SLOPE (3.5 HORIZONTAL TO 1 VERTICAL) ELEMENTS ELEMENT NUMBER CRITICAL NUMBER OF CYCLES Nc AVERAGE CYCLIC SHEAR STRESS FOR N c CYCLES avg (lb/ft²) AVERAGE CYCLIC SHEAR STRESS CAUSING 5% STRAIN IN Nc CYCLES fail (lb/ft²) MINIMUM FACTOR OF SAFETY WITH RESPECT TO 5% STRAIN FOR N c CYCLES fail/avg 8 6 177 520 2.94 9 6 196 760 3.88 16 6 323 520 1.61 17 6 322 940 2.92 19 5 352 1000 2.84 21 6 372 940 2.53 22 6 383 940 2.45 26 6 418 940 2.25 36 6 472 940 1.99 39 5 616 1180 1.92 42 5 677 1180 1.74 47 6 497 760 1.53 59 6 495 760 1.54 62 4 856 1240 1.45 65 5 908 1180 1.30 68 5 939 1180 1.26 72 6 461 760 1.65 75 5 870 1180 1.36 87 6 713 1095 1.54 88 6 797 1362 1.71 89 6 922 1629 1.77 90 6 972 2124 2.19 91 6 1063 2391 2.25 92 6 1092 2832 2.59 93 6 1088 2832 2.60 94 6 1068 2832 2.65 95 6 1036 2832 2.73 99 6 303 520 1.72 101 5 662 1000 1.51 121 5 769 1000 1.30 CPS/USAR CHAPTER 02 2.5-225 REV. 11, JANUARY 2005 TABLE 2.5-57 (Cont'd) (Q&R 241.16) LOCAL FACTORS OF SAFETY IN THE NATURAL SLOPE ELEMENT NUMBER CRITICAL NUMBER OF CYCLES Nc AVERAGE CYCLIC SHEAR STRESS FOR N c CYCLES avg (lb/ft²) AVERAGE CYCLIC SHEAR STRESS CAUSING 5% STRAIN IN Nc CYCLES fail (lb/ft²) MINIMUM FACTOR OF SAFETY WITH RESPECT TO 5% STRAIN FOR N c CYCLES fail/avg 98 8 366 495 1.35 100 5 531 1290 2.43 101 5 715 1500 2.10 102 5 923 1545 1.68 103 5 1041 1710 1.65 104 5 1138 1787 1.57 105 5 1256 1823 1.46 106 5 1326 1844 1.40 107 5 1386 1844 1.33 108 5 1413 1844 1.31 109 5 1331 1844 1.39 114 9 64 122 1.91 115 7 70 137 1.96 116 9 58 122 2.10 117 9 49 122 2.49 118 7 790 945 1.20 119 6 572 1155 2.02 120 5 675 1380 2.04 122 5 984 1562 1.59 123 5 1116 1727 1.55 124 5 1225 1806 1.47 125 5 1329 1832 1.38 126 5 1433 1865 1.30 127 5 1498 1868 1.25 128 5 1549 1868 1.21 129 5 1627 1868 1.15 CPS/USAR CHAPTER 02 2.5-226 REV. 11, JANUARY 2005 The factors of safety for the requested elements are given in the preceding table. The minimum factors of safety for sand elements were obtained by comparing the shear stresses required to cause single-amplitude shear strain of 5% (see Figure 2.5-413) with the equivalent shear stresses induced by the earthquake. It can be seen from the following table that the 5% strain occurs before the initial liquefaction starts. Therefore, the criterion of 5% strain is conservative, and is assures that the sand elements will not liquefy during the earthquake (Q&R 241.16).

TABLE 2.5-57 (Q&R 241.16) RESULTS OF STRESS-CONTROLLED CYCLIC TRIAXIAL (LIQUEFACTION) TESTS BORING OR TEST PIT NUMBER DEPTH OF SAMPLE ft. SOIL DESCRIPTION DEGREE OF COMPACTION MOLDING MOISTURE CONTENT,

% PRINCIPAL CONSOLIDATION STRESS RATIO kC LATERAL CONSOLIDATION PRESSURE 3C, psf CYCLE STRESS RATIO V/2 3c SKEMPTON'S PORE PRESSURE, PARAMETER, B 5% STRAIN1 10% STRAIN1 20% STRAIN1 INITIAL LIQUE-FACTION2 Interglacial Granular Soils:

TP-3 9-11 Bluish gray fine 75% 7.9 1.0 2000 0.21 0.97 66 70 92 92 to medium sand (Relative) 7.6 1.0 2000 0.39 1.0 8 11 20 9 with traces of 7.4 1.0 2000 0.61 0.95 3 5 10 5 silt and fine gravel H-23 49 Brown silty fine

--3 16.95 1.0 2000 0.38 0.99 25 90 200 200 sand H-31 14 Gray gravelly fine --4 8.65 1.0 2000 0.37 0.96 8 28 56 8 to coarse sand with some silt 1 Double Amplitude Axial Strain.

2 Initial liquefaction is defined as when the increase in pore pressure is equal to the effective confining pressure.

3 "Undisturbed" sample taken during the boring operations, dry density = 112.8 pcf.

4 "Undisturbed" sample taken during the boring operations, dry density = 125.6 pcf.

5 Initial (in-situ) moisture content.

CPS/USAR CHAPTER 02 2.5-227 REV. 11, JANUARY 2005 TABLE 2.5-58 SUMMARY OF LOCAL FACTORS OF SAFETY IN SUBMERGED DIKE ELEMENTS DUE TO A NEW MADRID TYPE EARTHQUAKE ELEMENT NUMBER CRITICAL NUMBER OF CYCLES N c AVERAGE CYCLIC SHEAR STRESS FOR N c CYCLES avg (lb/ft²) AVERAGE CYCLIC SHEAR STRESS CAUSING 5% STRAIN IN N c CYCLES fail (lb/ft²) MINIMUM FACTOR OF SAFETY WITH RESPECT TO 5% STRAIN FOR N c CYCLES fail/avg 1 20 11.6 175 15.09 2 20 19.3 175 9.07 4 20 29.1 175 6.01 6 20 32.2 175 5.44 11 30 27.6 80 2.90 12 10 53.5 313 5.85 14 10 71.1 313 4.40 15 10 75.9 313 4.12 16 10 75.6 313 4.14 21 30 31.7 80 2.52 22 10 78.1 313 4.01 23 10 97.6 432 4.43 24 10 111.1 432 3.89 25 10 117.7 432 3.67 26 10 117.5 432 3.68 33 10 129.8 432 3.33 35 10 164.9 610 3.70 36 10 164.7 610 3.70 43 20 57.3 320 5.59 45 20 98.3 320 3.26 47 13 185.7 610.8 3.29 49 13 207.4 763.2 3.68 50 13 205.4 763.2 3.72 61 13 203.2 610.8 3.01 64 10 281.2 820 2.92 66 10 306.8 820 2.67 68 10 312.0 820 2.63 81 10 426.1 820 1.92 85 10 453.9 820 1.81 87 10 461.2 1000 2.17 89 10 454.1 1000 2.20 CPS/USAR CHAPTER 02 2.5-228 REV. 11, JANUARY 2005 TABLE 2.5-59 SUMMARY OF LOCAL FACTORS OF SAFETY IN THE NATURAL SLOPE (3.5 HORIZONTAL TO 1 VERTICAL) ELEMENTS DUE TO A NEW MADRID TYPE EARTHQUAKE ELEMENT NUMBER CRITICAL NUMBER OF CYCLES N c AVERAGE CYCLIC SHEAR STRESS FOR N c CYCLES avg (lb/ft²) AVERAGE CYCLIC SHEAR STRESS CAUSING 5% STRAIN IN N c CYCLES fail (lb/ft²) MINIMUM FACTOR OF SAFETY WITH RESPECT TO 5% STRAIN FOR N c CYCLES fail/avg 8 20 97.7 320 3.28 9 10 128.7 660 5.13 16 20 163.9 320 1.95 17 6 257.8 950 3.69 19 6 261.4 950 3.63 21 6 269.1 950 3.53 22 7 235.1 910 3.87 26 6 345.0 950 2.75 36 6 398.1 950 2.39 39 4 521.8 1240 2.38 42 4 541.2 1240 2.29 47 6 426.6 770 1.81 59 6 432.2 770 1.78 62 4 709.2 1240 1.75 65 4 770.5 1240 1.61 68 4 760.7 1240 1.63 72 6 407.3 770 1.89 75 4 767.9 1240 1.62 87 20 394.0 648 1.65 88 20 433.1 806 1.86 89 18 521.2 1013.0 1.94 90 18 545.7 1320.7 2.42 91 18 592.5 1486.7 2.51 92 18 604.6 1760.9 2.91 93 18 599.6 1760.9 2.94 94 18 587.5 1760.9 3.00 95 18 572.6 1760.9 3.08 99 25 151.2 287 1.90 101 6 539.9 950 1.76 121 5 663.2 1000 1.51 CPS/USAR CHAPTER 02 2.5-229 REV. 11, JANUARY 2005 TABLE 2.5-60 ATTERBERG LIMITS DATA ULTIMATE HEAT SINK BORINGS BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT PLASTIC LIMIT

(%) PLASTICITY INDEX (%) H-2 663.9 SM/ML Salt Creek Alluvium 20.4 20.4 H-2 627.4 ML Illinoian Glacial Till 12.8 7.3 H-3 624.6 ML Illinoian Glacial Till 12.2 9.3 H-4 657.6 ML Illinoian Glacial Till 11.9 6.1 H-4 622.6 ML Illinoian Glacial Till 11.8 7.6 H-4 602.6 ML Illinoian Glacial Till 11.6 8.0 H-5 653.6 ML Illinoian Glacial Till 11.6 6.5 H-5 608.6 ML Illinoian Glacial Till 11.4 8.6 H-6 653.3 ML Illinoian Glacial Till 11.5 7.6 H-6 583.3 ML Illinoian Glacial Till 11.5 5.1 H-6 578.3 CL Illinoian Glacial Till 13.6 12.0 H-7 649.5 ML Illinoian Glacial Till 10.0 7.7 H-7 630.0 ML Illinoian Glacial Till 11.8 2.6 H-10 623.2 ML Illinoian Glacial Till 12.2 8.1 H-11 598.1 ML Illinoian Glacial Till 11.4 8.3 H-12 635.6 ML Illinoian Glacial Till 10.9 4.6 H-15 711.3 ML Wisconsinan Glacial Till 12.0 1.6 CPS/USAR TABLE 2.5-60 (Cont'd) CHAPTER 02 2.5-230 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT PLASTIC LIMIT

(%) PLASTICITY INDEX (%) H-18 645.5 ML Illinoian Glacial Till 11.5 6.9 H-20 701.8 ML Wisconsinan Glacial Till 13.5 11.6 H-25 638.7 ML Illinoian Glacial Till 11.3 5.6 H-31 689.1 ML Wisconsinan Glacial Till 13.1 16.4 H-31 680.1 CL Interglacial Zone 13.8 23.5 H-32 555.1 ML Pre-Illinoian Glacial Till 12.9 13.2 H-33 674.8 SM Salt Creek Alluvium N.P. N.P. H-36 541.2 ML Pre-Illinoian Glacial Till 12.8 11.9 H-3 654.6 ML Illinoian Glacial Till 11.5 6.5 H-5 634.1 ML Illinoian Glacial Till 10.8 5.5 H-8 658.3 SM Salt Creek Alluvium N.P. N.P. H-10 643.7 ML Illinoian Glacial Till 10.9 4.2 H-12 671.1 ML Salt Creek Alluvium 22.6 20.0 H-12 663.6 ML Illinoian Glacial Till 12.4 5.4 H-16 662.3 ML Illinoian Glacial Till 13.2 6.0 H-24 670.7 SM Salt Creek Alluvium N.P. N.P. H-24 627.7 ML Illinoian Glacial Till 12.3 6.1 H-25 658.2 ML Salt Creek Alluvium 11.3 5.8 CPS/USAR TABLE 2.5-60 (Cont'd) CHAPTER 02 2.5-231 REV. 11, JANUARY 2005 BORING NUMBER ELEVATION (ft-in.)

SOIL TYPE GEOLOGIC UNIT PLASTIC LIMIT

(%) PLASTICITY INDEX (%) H-29 682.7 ML Illinoian Glacial Till 14.8 11.8 H-32 620.1 ML Illinoian Glacial Till 11.9 8.6 H-33 662.8 ML Illinoian Glacial Till N.P. N.P. H-34 652.6 ML Illinoian Glacial Till 11.0 5.3 H-34 632.6 ML Illinoian Glacial Till 11.8 7.8 H-35 632.8 ML Illinoian Glacial Till 10.8 6.0 H-37 689.9 ML Interglacial Soil 20.1 22.6 CPS/USAR CHAPTER 02 2.5-232 REV. 11, JANUARY 2005 TABLE 2.5-61 MODIFIED MERCALLI INTENSITY (DAMAGE) SCALE OF 1931 I. Not felt except by a very few under especially favorable circumstances. (I, Rossi-Forel Scale) II. Felt only by a few persons at rest, especially on upper floors of buildings. Delicately suspended objects may swing. (I to II, Rossi-Forel Scale) III. Felt quite noticeably indoors, especially on upper floors of buildings, but many people do not recognize it as an earthquake. Standing motorcars may rock slightly. Vibration like passing of truck. Duration estimated. (III, Rossi-Forel Scale) IV. During the day, felt indoors by many, outdoors by few. At night, some awakened. Dishes, windows, doors disturbed; walls make creaking sound. Sensation like heavy truck striking building. Standing motorcars rock noticeably. (IV to V, Rossi-Forel Scale) V. Felt by nearly everyone, many awakened. Some dishes, windows, etc., broken; a few instances of cracked plaster; unstable objects overturned. Disturbances of trees, poles, and other tall objects sometimes noticed. Pendulum clocks may stop. (V to VI, Rossi Forel Scale) VI. Felt by all, many frightened and run outdoors. Some heavy furniture moved; a few instances of fallen plaster or damaged chimneys. Damage slight. (VI to VII, Rossi-Forel

Scale) VII. Everybody runs outdoors. Damage negligible in buildings of good design and construction; slight to moderate in well-built ordinary structures; considerable in poorly built or badly designed structures: some chimneys broken. Noticed by persons driving motorcars. (VIII, Rossi-Forel Scale) VIII. Damage slight in specially designed structures; considerable in ordinary substantial buildings, with partial collapse; great in poorly built structures. Panel walls thrown out of frame structures. Fall of chimneys, factory stacks, columns, monuments, walls. Heavy

furniture overturned. Sand and mud ejected in small amounts. Changes in well water. Persons driving motorcars disturbed. (VIII+ to IX-, Rossi-Forel Scale) IX. Damage considerable in specially designed structures: well designed frame structures thrown out of plumb; great in substantial buildings, with partial collapse. Buildings shifted off foundations. Ground cracked conspicuously. Underground pipes broken.

(IX+, Rossi-Forel Scale) X. Some well-built wooden structures destroyed: most masonry and frame structures destroyed with foundations; ground badly cracked. Rails bent. Landslides considerable from river banks and steep slopes. Shifted sand and mud. Water splashed (slopped) over banks. (X, Rossi-Forel Scale) XI. Few, if any, (masonry) structures remain standing. Bridges destroyed. Broad fissures in ground. Underground pipelines completely out of service. Earth slumps and land slips in soft ground. Rails bent greatly. XII. Damage total. Waves seen on ground surface. Lines of sight and level distorted. Objects thrown upward into the air.

CPS/USAR CHAPTER 02 2.5-233 REV. 11, JANUARY 2005 TABLE 2.5-62 SUMMARY OF THE CONSOLIDATION TEST DATA ATTERBERG LIMITS (percent) BORING NUMBER ELEVATION (feet) GEOLOGICAL MEMBER MOISTURE CONTENT (percent)

DRY DENSITY (pcf) INITIAL1 VOID RATIO SAMPLING METHOD PLASTIC LIMIT PLASTICITY INDEX COMPRESSION INDEX RE-COMPRESSION INDEX PRE-CONSOLIDATION 2 PRESSURE (psf) P-14 683.8 Interglacial Till 16.2 115 0.41 U 12.5 12.5 0.12 0.03 4,100 P-15 711.8 Wisconsinan Till 12.8 116 0.40 U 14.5 8.5 0.14 0.025 5,700 P-18 683.7 Illinoian Till 17.7 104 0.56 U NA NA 0.19 0.03 10,500 P-20 688.3 Interglacial Till 10.9 122 0.33 U 14.4 7.2 0.12 0.023 8,000 P-21 695.7 Interglacial Till 18.4 114 0.42 U 16.6 4.4 0.

07 0.011 >5,000 P-22 728.5 Loess 22.2 97 0.68 U 22.8 17.3 0.25 0.044 2,100 699.5 Interglacial Till 36.8 74 1.19 U NA NA 0.36 0.030 4,500 689.5 Interglacial Till 12.8 119 0.37 U 12.4 5.0 0.085 0.014 5,000 P-26 692.1 Interglacial Till 14.6 120 0.35 U 9.8 11.6 0.11 0.024 4,500 P-27 728.4 Wisconsinan Till 11.2 126 0.29 U 10.4 6.8 0.083 0.016 4,000 703.4 Interglacial Till 17.5 109 0.49 U 12.4 11.5 0.13 0.022 >4,000 P-29 706.5 Wisconsinan Till 7.5 121 0.34 U 16.8 9.2 0.11 0.014 3,800 667.0 Illinoian Till 17.7 110 0.47 P NA NA 0.094 0.017 17,000 627.0 Illinoian Till 9.9 134 0.21 P 11.0 5.7 0.089 0.016 22,000 617.0 Illinoian Till 10.1 136 0.19 U NA NA 0.11 0.017 19,500 577.0 Illinoian Till 7.6 141 0.15 P 11.7 35.0 0.082 0.008 22,000 522.0 Pre-Illinoian Till 14.9 116 0.40 P NA NA 0.15 0.036 23,000 517.0 Pre-Illinoian Till 16.9 116 0.40 P 15.7 9.7 0.16 0.036 14,000 CPS/USAR TABLE 2.5-62(Cont'd) CHAPTER 02 2.5-234 REV. 11, JANUARY 2005 ATTERBERG LIMITS (percent) BORING NUMBER ELEVATION (feet) GEOLOGICAL MEMBER MOISTURE CONTENT (percent)

DRY DENSITY (pcf) INITIAL1 VOID RATIO SAMPLING METHOD PLASTIC LIMIT PLASTICITY INDEX COMPRESSION INDEX RE-COMPRESSION INDEX PRE-CONSOLIDATION 2 PRESSURE (psf) P-32 667.4 Illinoian Till 8.4 136 0.19 C 11.3 10.8 -- 0.009 18,000 644.2 Illinoian Till 6.7 138 0.18 C 11.3 6.6 0.13 0.012 20,000 627.9 Illinoian Till 6.9 140 0.16 C 10.9 8.9 0.062 0.010 17,000 585.0 Illinoian Till 8.5 136 0.19 C 10.6 7.8 0.11 0.014 25,000 P-33 675.6 Illinoian Till 7.2 139 0.17 U NA NA 0.053 0.013 >20,000 P-36 639.2 Illinoian Till 6.5 140 0.16 C 10.7 6.5 0.059 0.011 18,000 613.7 Illinoian Till 8.6 135 0.20 C 11.9 9.0 0.13 0.014 20,000 599.2 Illinoian Till 7.9 138 0.18 C 11.8 7.0 0.090 0.009 25,000 559.2 Pre-Illinoian Till 6.8 137 0.18 C NA NA 0.094 0.010 21,000 538.7 Pre-Illinoian Till 16.3 116 0.4 C 17.8 25.5 0.24 0.020 23,000 P-38 655.6 Illinoian Till 7.7 138 0.18 C 9.2 6.0 0.11 0.007 16,000 638.4 Illinoian Till 6.4 139 0.17 C 9.7 5.8 0.095 0.009 17,000 634.9 Illinoian Till 6.9 139 0.17 C NA NA 0.053 0.010 >20,000 616.9 Illinoian Till 10.1 132 0.23 C NA NA 0.18 0.015 21,000 598.7 Illinoian Till 8.4 138 0.18 C 9.8 6.6 0.11 0.009 17,000 569.9 Lacustrine 11.7 126 0.29 U NA NA 0.11 0.020 24,000 567.6 Lacustrine 10.6 129 0.26 C 10.2 7.5 0.12 0.013 22,000 531.5 Pre-Illinoian Till 16.0 116 0.40 C 17.5 21.0 0.17 0.013 23,000 P-41 714.7 Wisconsinan Till 11.1 128 0.27 C NA NA 0.034 0.0053 5,000 679.7 Illinoian Till 7.5 138 0.18 C NA NA 0.031 0.014 15,500 CPS/USAR TABLE 2.5-62(Cont'd) CHAPTER 02 2.5-235 REV. 11, JANUARY 2005 LEGEND NA - Test result not available C inch core sampler P - Pitcher sampler U - Dames & Moore "U" sampler

Notes: 1. Initial void ratio computed by assuming specific gravity equal to 2.6. 2. Preconsolidation pressures determined by using A. Casagrande approach.

CPS/USAR CHAPTER 02 2.5-236 REV. 11, JANUARY 2005 TABLE 2.5-63 BEARING CAPACITY AND FACTOR OF SAFETY FOUNDATION AREA SEISMIC CATEGORY I STRUCTURE FOUNDATION ELEVATION (feet) GROSS APPLIED STATIC FOUNDATION PRESSURE (ksf) NET STATIC FOUNDATION PRESSURE*

(ksf) ULTIMATE BEARING CAPACITY (ksf) FACTOR OF SAFETY Containment X 702.0 6.5 4.8 90.1 18.8 Fuel Building X 702.0 6.5 4.8 110.4 23.0 Auxiliary X 697.5 6.5 4.5 91.8 20.4

Radwaste, Machine Shop, and Off-Gas Building X 692.0 4.8 2.4 121.2 50.5 Service Building 732.0 1.5 1.5 51.2 34.1 Diesel Generator and HVAC X 702.0 4.7 3.0 79.8 26.6 Control X 692 4.7 2.3 84.1 36.6 Turbine and Heater Bay 702.0 5.7 4.0 120.3 30.1 Circulating Water Screen House X 653.0 3.2 2.9 244.0 84.1 Ultimate Heat Sink

Outlet Structure X 669.0 0.6 .48 14.5 30.2 __________________________

• The net static foundation pressure equals the gross applied static foundation pressure minus the hydrostatic uplift pressure.

CPS/USAR CHAPTER 02 2.5-237 REV. 11, JANUARY 2005 TABLE 2.5-64 TYPICAL EXAMPLE OF CALCULATION OF FACTOR OF SAFETY STRESSES (lb/ft2) STRESSES IN DESCENDING ORDER (lb/ft

2) CYCLE +ve -ve +ve -veAVERAGE STRESS (lb/ft²) CUMULATIVE AVERAGE STRESS (lb/ft²) FAILURE (lb/ft²)

FACTOR OF SAFETY 1 113 204 169218194 194 620 3.20 2 132 218 150212181 188 520 2.77 3 149 162 149204177 184 470 2.55 4 150 130 132162147 175 430 2.45 5 80 106 118130124 165 400 2.42 6 169 212 113130122 158 370 2.34 7 48 65 111106109 151 360 2.38 8 111 130 106106106 145 350 2.40 9 118 28 106103105 141 340 2.41 10 106 103 857781 135 330 2.44 11 30 38 806573 129 325 2.52 12 71 77 715362 123 320 2.60 13 14 106 525252 118 315 2.67 14 52 31 483843 113 310 2.74 15 106 53 303131 107 305 2.85 16 85 52 142821 102 300 2.94 CPS/USAR CHAPTER 02 2.5-238 REV. 11, JANUARY 2005 TABLE 2.5-65 SEISMIC SOURCE PARAMETERS SOURCE ACTIVITY RATE (MEAN NUMBER OF EARTHQUAKES/YEAR)

HISTORICAL MAXIMUM INTENSITY UPPER BOUND INTENSITY 1. Illinois Basin Seismogenic Region 0.20 VII VIII 2. Ste. Genevieve Seismogenic Region 0.02 VI VII 3. St. Francois Mountains Seismogenic Region 0.03 VI-VII VII-VIII 4. Chester Dupo Seismogenic Region 0.27 VII VIII 5. Wabash Valley Seismogenic Region 0.17 VII VIII 6. Western Kentucky Fault Zone Seismotectonic Region 0.03 V VI

7. Iowa-Minnesota Seismogenic Region 0.01 V VI 8. Missouri Random Seismogenic Region 0.02 V VI 9. Michigan Basin Seismogenic Region 0.02 VI VII 10. Eastern Interior Arch System Seismogenic Region 0.04 VII VIII 11. Anna Seismogenic Region 0.25 VII-VIII VIII-IX 12. New Madrid Seismogenic Region 0.33 XII XII CPS/USAR CHAPTER 02 2.5-239 REV. 11, JANUARY 2005 TABLE 2.5-66 RECOMPRESSION INDEX FOR SETTLEMENT ANALYSIS Cr/ (1+eo) ELEVATIONS RELOADING BELOW Po AND REBOUND RELOADING ABOVE Po BUT BELOW P C 660 to 680 ft 0.005 0.0110 627 to 660 ft 0.005 0.0083 614 to 627 ft 0.005 0.0142 575 to 615 ft 0.005 0.0091 560 to 575 ft 0.005 0.0165 540 to 560 ft 0.005 0.0117 522 to 540 ft 0.005 0.0171 500 to 522 ft 0.005 0.0257 _________________________ NOTES:

1. Po = In situ overburden pressure

2. PC = Overconsolidation pressure

3. Cr = Recompression index

4. eo = Initial void ratio CPS/USAR CHAPTER 02 2.5-240 REV. 11, JANUARY 2005 TABLE 2.5-67 COMPARISON OF CALCULATED AND MEASURED SETTLEMENTS BUILDING SETTLEMENT MONUMENT* CALCULATED FINAL SETTLEMENT (in.) MEASURED*** SETTLEMENT (in.) DATE OF FIRST MEASUREMENT DATE OF LAST MEASUREMENT Containment C1A** 1.72 -0.096 Jan. 1981 Sept. 1983 C2 1.72 0.264 May 1978 Jan. 1984 C3 1.64 0.108 May 1978 Sept. 1983 C4 1.79 0.504 May 1978 May 1982 C5 1.75 0.384 May 1978 Jan. 1984 C7 1.59 0.384 May 1978 Jan. 1984 C8A** 1.72 -0.048 Jan. 1981 Jan. 1984 C9 1.59 0.420 May 1978 May 1982 C10 1.61 0.360 May 1978 Jan. 1984 C11 1.37 0.492 May 1978 Jan. 1982 Turbine T1A** 1.45 -0.048 Sept. 1980 Jan. 1984 T2 1.44 0.228 May 1978 Jan. 1984 T3A** 1.40 0.096 Sept. 1980 Jan. 1984 T4A** 1.36 -0.060 Nov. 1980 Jan. 1984 Diesel DIAB** 1.22 -0.036 Nov. 1980 Jan. 1984 Generator D2 1.47 0.360 May 1978 Jan. 1984 & Control D3 1.18 0.252 May 1978 Jan. 1984 D4A** 1.42 -0.072 Mar. 1981 Jan. 1984 D5 1.11 0.240 May 1978 Jan. 1984 D6 1.01 0.444 May 1978 May 1983 D7 1.26 0.180 May 1978 Jan. 1984 Radwaste R1A** 1.15 -0.156 Mar. 1980 Jan. 1984 Off-Gas & R2B** 1.05 -0.060 Nov. 1980 Jan. 1984 Machine Shop R3 1.20 0.024 May 1978 Jan. 1984 R4A** 0.97 -0.168 May 1980 Jan. 1984 R5A** 1.07 -0.168 Sept. 1981 Jan. 1984 __________________________

• See Figure 2.5-382 for locations of settlement monuments. ** D1 and D1A were replaced by D1B, R1 by R1A, R2 and R2A by R2B, R4 and R4A, R5 by R5A, C1 by C1A, C8 by C8A, T1 by T1A, T3 by T3A, T4 by T4A, and D4 by D4A. *** Representing difference between first and last readings. Negative signs indicate heave.

CPS/USAR CHAPTER 02 2.5-241 REV. 11, JANUARY 2005 TABLE 2.5-68 (Q&R 241.16) RESULTS OF STRESS-CONTROLLED CYCLIC TRIAXIAL (LIQUEFACTION) TESTS NUMBER OF CYCLES TO CAUSE BORING OR TEST PIT NUMBER DEPTH OF SAMPLE ft. SOIL DESCRIPTION DEGREE OF COMPACTION MOLDING MOISTURE CONTENT,

% PRINCIPAL CONSOLIDATION STRESS RATIO Kc LATERAL CONSOLIDATION PRESSURE 3c, psf CYCLE STRESS RATIO v/2 3c SKEMPTON'S PORE PRESSURE, PARAMETER, B 5% STRAIN1 10% STRAIN1 20% STRAIN1 INITIAL LIQUEF-ACTION2 Interglacial Granular Soils:

TP-3 9-11 Bluish gray fine 75% 7.9 1.0 2000 0.21 0.97 66 70 92 92

to medium sand (Relative) 7.6 1.0 2000 0.39 1.0 8 11 20 9 with traces of 7.4 1.0 2000 0.61 0.95 3 5 10 5 silt and fine gravel H-23 49 Brown silty fine

-3 16.95 1.0 2000 0.38 0.99 25 90 200 200 sand H-31 14 Gray gravelly fine -4 8.65 1.0 2000 0.37 0.96 8 28 56 8 to coarse sand with some silt 1 Double Amplitude Axial Strain.

2 Initial liquefaction is defined as when the increase in pore pressure is equal to the effective confining pressure.

3 "Undisturbed" sample taken during the boring operations, dry density = 112.8 pcf.

4 "Undisturbed" sample taken during the boring operations, dry density = 125.6 pcf.

5 Initial (in-situ) moisture content.

CPS/USAR CHAPTER 02 A2.5-1 REV. 11, JANUARY 2005 ATTACHMENT A2.5 ULTIMATE HEAT SINK LIQUEFACTION ANALYSIS

CPS/USAR CHAPTER 02 A2.5-2 REV. 11, JANUARY 2005 ATTACHMENT A2.5 - ULTIMATE HEAT SINK LIQUEFACTION ANALYSIS A2.5 ULTIMATE HEAT SINK LIQUEFACTION ANALYSIS A2.5.1 Summary Four failure modes were postulated which could infringe upon the capability of the Clinton Power Station (CPS) proposed ultimate heat sink (UHS) to perform its function. These four postulated failure modes are: a. loss of cooling water inventory due to its displacement by alluvial flow slides into the UHS; b. loss of the service water system due to blockage of the service water pump intakes from flow blocking or entering the intake structure; c. loss of the UHS circulation pattern due to local slides producing dikes or dams across the circulation channel; and d. loss of UHS water because the UHS dam or its flanks are breached by the combination of seismic loadings, liquefaction, and washout. Field exploration programs, laboratory testing programs, evaluation of soil instability through empirical and analytical methods, and evaluation of the consequences of postulated soil instability have led to the conclusion that the capability of the ultimate heat sink cannot be compromised by those failure modes. This attachment summarizes the investigations used to arrive at the above conclusion.

A2.5.2 Ultimate Heat Sink Description The ultimate heat sink is a submerged pond lying in the bottom of the cooling lake (Lake Clinton). A compacted earth dam lying across the lower portion of the North Fork stream valley retains the pond, and the required storage capacity was developed by excavating the valley alluvium to provide 1067 acre-feet storage capacity. In the event of loss of the cooling lake, recirculation within the pond will be guided by a baffle dike to gain the effective surface area required for heat dissipation. The dam and the baffle dike were constructed from cohesive soils available in the site area. These earth structures were founded on the Illinoian glacial till of the unaltered Glasford Formation which underlies the alluvial deposits presently in the stream valley. Figure A2.5-1 presents a plan of the ultimate heat sink. A2.5.3 Design Bases The ultimate heat sink will provide sufficient water volume and cooling capability for the station for 30 days with no water makeup. Subsection 9.2.5.3 presents the safety evaluation for the UHS. The ultimate heat sink will also provide a minimum of 900,000 gallons of water for fire protection, if required.

CPS/USAR CHAPTER 02 A2.5-3 REV. 11, JANUARY 2005 In addition to the storage requirements for cooling purposes and fire water supply, storage capacity is provided to accept some sedimentation. Sediment accumulation within the heat sink will be periodically measured, and in the event that accumulation exceeds the capacity provided for sediment storage, dredging will be performed. Significant design bases used in the determination of the minimum UHS volume include the following: a. sediment inflow from liquefaction - 221 acre-feet; b. required fire water storage capacity - 3 acre-feet; c. minimum cooling capacity requirement of UHS to meet 95

° F shutdown service water inlet temperature - 590 acre-feet; d. an uninterrupted flow path must be maintained at all times between the UHS inlet structure and the shutdown service water pumps; and e. maximum sedimentation due to a 100-year flood - 35 acre-feet. A2.5.4 Geotechnical Investigation A2.5.4.1 Introduction An evaluation of the soil slopes adjacent to the ultimate heat sink was conducted to determine the horizontal and vertical distribuition, relative denseness, and geometry of these potentially liquefiable alluvial soils. The field investigation portion of the program consisted of fourteen BH-series borings and twelve PH-series auger probes. The location of the borings and the probes are shown in Figure A2.5-2. To determine the consistency of the subsurface strata, borings were drilled with truck mounted rotary wash equipment using 4-inch continuous-flight augers, with samples being extracted by utilizing a standard 2-inch split-spoon sampler. A graphic representation of soils encountered in the borings, including standard penetration test data and sampling information, is presented in Figures A2.5-3 through A2.5-28. The method of classifying the soils is described in Figure 2.5-355.

A key to the sample symbols and samplings information presented on the logs of the borings is shown in Figure A2.5-29. The laboratory portion of the geotechnical investigation consisted of a program to identify the physical characteristics of the soils. Testing of the samples consisted of determining the Atterberg limits, moisture content, particle size, and unit weight determination. The results of these tests are shown in the boring logs presented as Figures A2.5-3 through A2.5-28. In areas where the alluvium is present, the generalized profile consists of a cohesive deposit underlain by a granular deposit which is in turn underlain by competent till materials of the unaltered Glasford Formation. A study of the heat sink's periphery reveals that certain sections of the bordering soil deposits may be characterized as possessing common geometrical configuration and material CPS/USAR CHAPTER 02 A2.5-4 REV. 11, JANUARY 2005 composition. Therefore, the heat sink's periphery has been divided into subsections identified by letters A through F in Figure A2.5-1 and designated as follows: a. Subsection I - Till Slopes from A to B; b. Subsection II - Till Slopes with Fringe of Alluvium from B to C; c. Subsection III - Retaining Fill from C to D; d. Subsection IV - Northwest Corner Alluvium Pocket from D to E;

e. Subsection V - Outwash Deposits from E to F; and

f. Subsection VI - Upstream Bed Alluvium from F to A. A2.5.4.2 Subsection I - Till Slopes from A to B As shown in Figure A2.5-1, the eastern end of the heat sink is bordered by till slopes which were graded to ensure stability under all loading conditions. Both the service pump structure and the SSWS outlet structure lie within this subsection, and the structures were founded on stable Illinoian glacial till of the unaltered Glasford Formation. The glacial tills which form the slopes are not susceptible to liquefaction, and during construction these slopes were graded to a uniform slope of 5:1 (horizontal to vertical). A2.5.4.3 Subsection II - Till Slopes with Fringe of Alluvium from B to C The south side of the heat sink is formed by alluvial soil deposits which lie at the base of natural till slopes which rise above the heat sink. The horizontal extent of the alluvium is shown by the crosshatched area in Figure A2.5-1. The quantity of alluvium postulated to slide into the heat sink from this fringe has been determined by assuming a 30:1 slope (horizontal to vertical), to

be approximately 90 acre-feet. A2.5.4.4 Subsection III - Compacted Earth Retaining Fill from C to D A compacted earth dam was constructed from cohesive materials not susceptible to liquefaction. Soil borings located in Figure A2.5-2 in the area of the dam have been used to determine the alluvium-Illinoian till interface. Section 11-11 (for plant location, see Figure A2.5-

2) in Figure A2.5-30 shows the existing alluvium-till interface; it also shows that existing alluvial material was removed and that the dam was constructed on Illinoian glacial till. The retaining fill spans the valley, and its ends intersect Illinoian glacial till at elevation 675 feet on each end. No material from Subsection III will move and displace the ultimate heat sink storage volume. A2.5.4.5 Subsection IV - Northwest Corner Alluvium Pocket from D to E A pocket of alluvium lies in the northwest corner of the heat sink. Exploratory borings, BH and PH series, were drilled in the area to confirm the contact between the alluvium and the Illinoian till. These borings are located in the plan in Figure A2.5-2. Since this alluvium is potentially susceptible to liquefaction, liquefaction is postulated by assuming a 30:1 slope (horizontal to vertical). It is postulated that 80 acre-feet of material will slide into the heat sink from this section.

CPS/USAR CHAPTER 02 A2.5-5 REV. 11, JANUARY 2005 A2.5.4.6 Subsection V - Outwash Deposits from E to F The valley of the North Fork of Salt Creek was formed by melt-water from the Wisconsinan ice sheet (Woodfordian Substage) eroding through Wisconsinan till of the Wedron Formation and down to the lllinoian till of the unaltered Glasford Formation. Subsequent to the downcutting action, outwash material consisting of gravel, sand, and silt was deposited on the Illinoian till within the valley, partially filling it. The outwash material has been identified as part of the Henry Formation and generally varies from clean, well graded sands with some fine gravel to silty sands and silts (Reference 3). The clean sands generally form the basal part of the outwash and the silty sands generally form the upper part. With the retreat of the ice sheet, the stream eroded the present channel in the outwash material leaving the terrace that is north of the heat sink. The outwash material that forms the terrace is capped in places by colluvial material which has been deposited from late Wisconsinan time to the present. The colluvial material is a silty clay with a trace of sand and fine gravel. The material is generally derived from erosion of the till uplands. BH series borings and PH series probes used to evaluate the outwash material in the terrace deposits are located in the plan in Figure A2.5-2. The terrace deposits consist of medium stiff clayey silt material capping the outwash materials. The outwash materials generally consist of two layers. The top layer is composed of approximately 4 feet of loose silt and sand deposits with an average blow count of N = 6. The silty sand is underlain by a layer, approximately 11 feet thick, of well graded, dense sand which directly overlies the glacial till. The outwash materials surface has a relatively flat terrain with an average rise of 10 feet in 900 feet. Slopes bordering the northern perimeter of the ultimate heat sink were examined by drawing five cross sections, Section 7-7 through 11-11, which are shown in Figures A2.5-31 and A2.5-30, respectively, and are located in Figure A2.5-2. These sections show the subsurface conditions in the area. In order to establish a conservative final slope for the evaluation of the potential for liquefaction of alluvial soils, the mechanism required to produce significant deformation and movement of the existing slopes must be considered. During the earthquake, before the liquefaction of the alluvium occurs, the soil in the existing slopes is subjected to varying shear stresses which in turn produce excess pore pressures. However, as the pore pressures build up in the granular alluvium, its ability to propagate seismic shear waves would be progressively diminished. Further, when the soil is liquefied prior to any movement, additional development of shear stresses would greatly be reduced. Finally, during the movement of soil after liquefaction there would be a further reduction in the pore pressure buildup through release of pore water. On this basis, the movement of liquefied soils is restricted to a finite slope dependent on site conditions. This evaluation uses a conservative slope of 30:1 (horizontal to vertical). This slope is based on past case histories (Reference 1) and collaborated by the testimony of J. Greeves of the NRC staff (Reference 2). The flow slide was postulated to determine the final

configuration of this area after SSE. By applying the 30:1 slope criterion, and balancing the soil volumes displaced above the final slope with the volume of the soil filling the original cross section below the final grade, the resultant cross section is established. The initial and final sections which were investigated for CPS/USAR CHAPTER 02 A2.5-6 REV. 11, JANUARY 2005 this analysis are drawn to natural scale and shown in Figure A2.5-31. The following observations are made in regard to the resultant configurations: a. Major portions of loose sand likely to liquefy during a SSE (identified in the cross section as SM layer) lies at a slope less than 30:1 (horizontal to vertical) and hence would not slide into the heat sink. b. Major portion of material carried into the heat sink with the postulated slide consists of plastic, silty clay, which due to its physical characteristics is unlikely to liquefy and flow unless taken for a ride along with underlying liquefied loose silty sand (SM) layer. On this basis the final cross sections indicated that the changed geometry of the original cut slopes will not impede the circulation of the cooling water. At Section 8-8, the toe of the postulated slope is 280 feet from the baffle dike. Moreover, the increase in the surface area as a result of a slide will somewhat make up for the restricted flow and add to the amount of water circulating in that section. The total material assumed to flow into the UHS is 50 acre-feet. The quantity is based on the area shown in Section 8-8 which shows the greatest amount of material that will flow in the heat sink in this section. A2.5.4.7 Subsection VI - Upstream Bed Alluvium The movement of the upstream bed alluvium is discussed in Reference 1. No movement of material is anticipated into the ultimate heat sink considering the viscous properties of the liquefied soil mass. Under extremely conservative conditions, less than 1 acre-foot is postulated to flow into the UHS. The postulated flow, in the form of a block, is approximately 20 feet into the UHS leaving it more than 1100 feet away from the screen house. A2.5.5 Conclusions Four failure modes were postulated which could infringe upon the capability of the CPS proposed ultimate heat sink (UHS) to perform its function. These four postulated failure modes

are: a. (Loss of cooling water inventory due to its displacement by alluvial flow slides into the UHS.) Postulated upper bound slides were used to evaluate potential heat sink capacity losses. The resulting losses are associated with the following perimeter subsection identified by letters A through F in Figure A2.5-1. Subsection I - Till Slopes from A to B; 0 acre-foot Subsection II - Till Slopes with Fringe of Alluvium from B to C; 90 acre-feet Subsection III - Retaining Fill from C to B; 0 acrefoot Subsection IV - Northwest Corner Alluvium Pocket from D to E; 80 acre-feet Subsection V - Outwash Deposits from E to F; 50 acre-feet CPS/USAR CHAPTER 02 A2.5-7 REV. 11, JANUARY 2005 Subsection VI - Upstream Bed Alluvium from F to A; 1 acre-foot. The evaluation for displacement of water by alluvium does not take credit for the change in pond configuration which will occur after the postulated slide has taken place. Sloughing of the north and south sides of the pond could, in effect, increase the pond's cooling capability by increasing surface area. For the ultimate heat sink capacity requirements see USAR Subsections 2.4.11.6 and 9.2.5. b. Loss of the shutdown service water system due to blockage of the shutdown service water pump intakes cannot occur because unstable soils do not lie at an elevation above the pump intakes in the area of the pump structure. The closest alluvium higher than the pump structure's intakes occurs at point A shown in Figure A2.5-1. This point lies approximately 700 feet from both the pump structure and the baffle dike. The analyses presented in "Potential for Alluvium in North Fork of Salt Creek to Flow into Heat Sink and Prevent Cooling Water Flow" (Reference 1), an unpublished report which has been submitted to the NRC Staff, establishes that the upstream alluvium at point A cannot block flow into the pump intakes. c. Loss of the UHS circulation pattern due to local slides has been evaluated by considering the final configuration of the postulated slides. Subsection II of the heat sink border is approximately 3600 feet long and along this distance 90 acre-feet is postulated to slide into the sink. The 90 acre-feet of material is uniformally distributed along Subsection II, and its cross-sectional area (perpendicular to the flow path) is approximately one-fifth of the area of the initial channel between the border and the baffle dike. Since material is available to block only one-fifth of the flow path along Subsection II, and since this material lies in space which would be vacated by soil and become heat sink capacity in the event of a slide, blockage of the circulation pattern along Subsection II is not considered to be a credible postulation. Cross section 8-8 presented in Figure A2.5-31 (location identified in Figure A2.5-2) is located at the highest terrain along the north edge. The configuration of the heat sink's cross section following postulated sliding is shown on Section 8-8 where the cross-hatched area indicates material that has slid into the heat sink.

The toe of the slide material lies 280 feet from the toe of the baffle dike. At the east and west end of the UHS, no change in the surface area is expected due to the stable till slopes and compacted dike respectively. d. Loss of ultimate heat sink water through the earth retaining fill cannot occur because: 1. The fill is compacted cohesive soil not susceptible to liquefaction.

2. The fill was founded on Illinoian glacial till and it abuts Illinoian glacial till which forms the valley walls. 3. The fill is protected with soil cement to withstand extreme event flow velocities.

CPS/USAR CHAPTER 02 A2.5-8 REV. 11, JANUARY 2005 4. The fill is designed to be stable under seismic loading. It is, therefore, concluded that the ultimate heat sink will provide sufficient water volume and cooling capability for the station for 30 days with no water makeup.

A2.5.6 References 1. I. M. Idriss and D. M. Hendron, "Potential for Alluvium in North Fork of Salt Creek to Flow into Heat Sink and Prevent Cooling Water Flow, Task I," Technical Report, Woodward-Clyde Consultants, April 30, 1975. 2. Transcripts of the ACRS 180th General Meeting, Washington, D.C., pp. 192-230, April 4, 1975. 3. J. P. Kepton, 1975 Illinois State Geological Survey, telephone communication, R. W. Wagner, April 21, 1975.

CPS/USAR CHAPTER 02 B2.5-1 REV. 11, JANUARY 2005 ATTACHMENT B2.5 SAND LENSES UNDER CATEGORY I STRUCTURES

CPS/USAR CHAPTER 02 B2.5-2 REV. 11, JANUARY 2005 ATTACHMENT B2.5 - SAND LENSES UNDER CATEGORY I STRUCTURES B2.5 SAND LENSES UNDER CATEGORY I STRUCTURES B2.5.1 Summary Postulating the liquefaction of sand lenses under Category I structures the following information was requested at the PSAR stage and is presented in this attachment. a. Evaluate the total and differential settlement caused by liquefaction of the sand lenses under earthquake loading. Consider both the regional SSE and the long duration event located in the Wabash Valley. Discuss the effect of this settlement on safety-related Category I foundations. State or reference the structural design criteria for settlement and differential settlement, and show that these criteria will not be exceeded. b. Describe the remedial treatment proposed for assuring that liquefaction of these sand lenses under earthquake loading will be presented. Soil borings, laboratory test data, relative densities based on standard penetration test values, grain size analyses, material type, in situ dry densities, and confining pressures have led to the conclusion that the sand lenses do not have characteristics of sand deposits capable for liquefaction under the postulated earthquake loadings; therefore, no remedial measures are needed and no settlements anticipated. B2.5.2 Subsurface Conditions The subsurface conditions in the area of the station mat foundations have been investigated with a series of borings; the boring logs are presented in Figures 2.5-19 through 2.5-73. Those borings which penetrate the area within the outline of the station's mat foundations are located in Figure B2.5-1. The soil under the station site and above approximately elevation 683 feet was removed and the excavation filled with Type B granular material recompacted to 85% relative density as determined by ASTM D-2049.

The Illinoian glacial till beneath the station site was located between elevations 575 and 686 feet, and the boring logs showed scattered sand lenses within the upper 30 feet of the Illinoian till. The strata in which the sand lenses lie is a glacial outwash area formed during the Illinoian glacial retreat. In outwash zones there is a high probability for the formation of scattered sand and gravel lenses and layers during fluctuation of the river flowing from the glacier. The geological process involved in the formation of the top 30 feet of the Illinoian till accounts for the apparent random scattering of granular material within the upper strata of the till. Figure B2.5-2 presents a plot of all granular material that has been penetrated by borings located within the outline of the station's mat foundation. Although the granular material cannot be definitely identified as lying in specific pockets and lenses, for purposes of discussion the granular material has been identified as lying in six distinct lenses as discussed herein.

CPS/USAR CHAPTER 02 B2.5-3 REV. 11, JANUARY 2005 B2.5.2.1 Additional Geotechnical Investigation In order to verify the assumptions made in determining the characteristic s of sandy soil forming these lenses, additional geotechnical investigation consisting of seven soil borings was performed. The seven borings (P-33A, P-50 through P-55) are located in Figure B2.5-1 and the subsurface conditions are shown on the boring logs in Figures B2.5-3 through B2.5-9. The program consisted of four primary borings (P-33A, P-50, P-52, and P-54) in which continuous sampling and standard penetration test values were taken in the sand lenses by means of a split spoon. The secondary borings lo cated by P-51, P-53, and P-55, were primarily intended to retrieve sand samples for laboratory testing. The samples in the secondary holes were obtained by either pitcher sampler or Dames & Moore 'U' sampler. Modified Osterberg sampler was also used to try to obtain undisturbed samples, but due to dense sand deposits no samples could be recovered. The laboratory testing program consisted of particle size analysis including hydrometer, Atterberg limit, and dry densities on relatively undisturbed samples. The results of the tests are shown on boring logs and Figures B2.5-10 through B2.5-15. B2.5.3 Soil Characteristics Influencing Liquefaction It is an established fact that liquefaction potential of soil deposits due to earthquake motion depends on characteristics of the soil, the initial stresses acting on the soil, and the characteristics of the earthquake involved (Reference 1). Significant factors include: a. The relative density Relative density is the most important physical characteristic that determines the liquefaction potential of a soil. The higher the relative density, the less susceptible the soil is to liquefaction. b. The soil type Fine sands and fine to medium sands tend to liquefy more easily than do coarse sands, gravelly soils, fine silts or clays. There is some evidence to show the poorly graded materials are more susceptible to liquefaction than well graded

materials. c. The initial confining pressure The liquefaction potential of a soil is reduced by an increase in confining pressure. State-of-the-art evaluation of soil characteristics for seismic response analyses (Reference 2) states, "From field observations it has generally been concluded by a number of investigators that even in a saturated sand deposit below a depth of 50 to 60 feet, sands are not likely to liquefy. These depths are in general agreement with Kishida (Reference 3) who states that a saturated sandy soil is not liquefiable if the value of the effective overburden pressure

exceeds 2 kg/cm 2 (kg/cm2 ~ 60 ft of soil below water table ~ 4.1 kips/sq ft)." Characteristics of earthquakes for this site are defined in Subsection 2.5.2.6.

CPS/USAR CHAPTER 02 B2.5-4 REV. 11, JANUARY 2005 B2.5.4 Liquefaction Potential of Sand Lenses The sand lenses under the main plant foundation are characterized by appreciable fines (passing U.S. sieve No. 200) ranging from 20% to as much as 55%. The high blow counts in the lenses are substantiated by high values of dry densities obtained from the extracted samples on relatively undisturbed samples in the laboratory. Shown in Figure B2.5-2 are the sand lenses under the main plant foundations. The sand lenses have been plotted to show the elevations at which these have been found and the corresponding boring numbers. A study of the sandy soil under the plant was performed which utilized the following parameters in delineating sand lenses: a. soil description, b. soil classification, c. penetration values, and

d. elevation at which the lenses exist. By studying these four parameters, the sand beneath the station can be classified into six distinct lenses. The characteristics of these lenses have been tabulated in Table B2.5-1. Sand Lens No. 1 This sand lens is located by Boring P-32 between elevations 675 feet and 677.5 feet. The lens has a Dames & Moore sampler penetration value of 10. The material is described as "Brown and gray, medium to coarse sand with trace of fine sand and silt," and is classified as SP-SW material. This lens was investigated after the foundation excavation and was completely removed. Sand Lens No. 2 A lens of SM material described as "Gray silty fine to coarse sand with a trace of fine gravel" is evident in Boring P-34. The extent of this layer is from elevation 677 feet to 679.5 feet. The standard penetration test N value exceeds 200, indicating extremely dense consistency.

As shown in Table B2.5-1, this sand has a relative density exceeding 95% and therefore is not susceptible to liquefaction. Sand Lens No. 3 This lens underlies most of the main plant site and appears in Borings P-10, P-14, P-30, P-33, P-34, P-35, P-36, P-37, P-41, P-33A, and P-52. This lens extends from elevation 652 feet to 665.5 feet. The soil description generalized from the 15 samples is "Gray fine to coarse sand with some silt and fine gravel" and can be classified as SM-SP-SW material. A major portion of sand under the main plant structures is included in this lens, and has a mean elevation of 658 feet. In order to determine the representative relative density of the lens, a statistical analysis was performed on the various penetration values and the result of this study has been summarized in Table B2.5-1.

CPS/USAR CHAPTER 02 B2.5-5 REV. 11, JANUARY 2005 Statistical Property Evaluation of Sand Lens To accurately represent the soil properties in the large sand lens, shown in Figure B2.5-2, a statistical analysis was performed. This type of analysis allows for variations in the testing procedure and will yield a probabilistic range of values. The first step in this procedure is to reduce the field test data. Nineteen standard penetrations and Dames & Moore samples were taken in a sand lens, which appears, by soil description and elevation similarity, to be the same lens. The corrected standard penetration N-Values were used to statistically compute a mean value to be used in relative density calculations. The analysis was performed assuming a normal distribution for the N-Values within the layer. There is a 97.5% probability that the five mean values will be greater than 68 blows per foot. This analysis was performed with the methods shown in Reference 4. Utilizing the unit weights for tills and the distances indicated on the boring logs, the vertical effective overburden pressure was found for each of the 19 samples. A statistical analysis of these two values was performed to achieve a confident range of values to use in the relative density calculations. A range of the arithmetic mean

+/- one standard deviation of the samples was used and this yielded a range of 5.68 to 6.40 ksf. With the N value and vertical effective overburden pressure, the Gibbs & Holtz relationship (Reference 5) for relative density was used. A range of values, based on the range of effective overburden pressures, was achieved. This range was from 87% to 92% relative density, and an average relative density of 90% was assigned to Lens 3. Sands with a relative density of 90% are unlikely to liquefy under the given confining pressures and anticipated loading (Reference 3). Sand Lens No. 4 Boring P-10 indicated a lens of "Gray fine to coarse sand with trace silt" (SW) from elevation 672.5 feet to 669 feet. The split spoon sample showed a standard penetration test value of 64 which indicates a material of very dense consistency.

As shown in Table B2.5-1, this sand lens has a relative density exceeding 90%, and therefore is not susceptible to liquefaction. Sand Lens No. 5 This sand lens is located north of the plant site and was originally shown in Borings P-33 and P-41. In order to better define this lens, three primary borings (P-33A, P-50, and P-52) were drilled at the locations shown in Figure B2.5-1, and continuous sampling of the sand lens was done with 'N' values taken at 1.5 feet intervals. As shown by the boring logs, P-33A and P-50, the sand lens does not extend under the station mat at the northeast section of the plant. Boring P-52 shows very dense sand from elevation 662.5 feet to 675.0 feet with an average standard penetration value of 71 (average of seven values). The sand in this lens is mostly medium to coarse sand with little gravel and silt; it is identified as well graded material (SW-SM-SP). Based on the standard penetration values of 71 and using the criteria proposed by Gibbs & Holtz, the relative density of the lens is approximately 95%. The characteristics of liquefaction in this sand lens deposit are absent; hence, it is not susceptible to liquefaction.

CPS/USAR CHAPTER 02 B2.5-6 REV. 11, JANUARY 2005 Sand Lens No. 6 The grain size analysis shown in Figure B2.5-16 indicates fines up to 55% in the upper part of the lens and, as indicated by Reference 6, it is not susceptible to liquefaction. The low blow counts of 23 and 13 are inconsistent with the dry density of 133 pcf obtained from the laboratory testing of the sample and, therefore, the criteria of relative density is not applicable. This lens, previously shown only in Boring P-38 was examined by the primary Boring P-54 taken in the vicinity as shown in Figure B2.5-1. Based on the sieve analysis data presented in Figure B2.5-16, the material can be classified as ML-SM and has a percent passing the U.S. sieve No.

200 up to 45% in the lower portion of the lens. The soil Boring P-54 taken in the vicinity of this sand lens has standard penetration values N of 81 and 50. The average standard penetration N value of this sand lens is 50 and has a relative density of 85%. Based on this information, we conclude that this sand lens is not susceptible to liquefaction. B2.5.5 Conclusions Sand lenses under the plant site have been examined for: a. relative density,

b. soil type, and c. initial confining pressure. From the numerous studies conducted on sands both in the laboratory and field, these three are the principal soil characteristics effecting liquefaction of sand deposits under earthquake loading. The properties of sand under the foundation mat have been examined for these characteristics and, in all cases, it has been found that the lenses will not liquefy under the earthquake loading. The consistency of the lenses was based on standard penetration test values; all the lenses (except Lens 1) have a relative density greater than 85%. The dense nature of the sands is borne out by a limited number of in situ dry density values of sands (P-6, P-13, P-15, P-20, P-26, P-43, H-16) where dry densities are found to be more than 122 pcf which are greater than 95% relative density of Sangamonian sands in the area. During the excavation and subgrade testing, Lens 1 was completely removed. The material of the sand lenses is usually medium to coarse sands, with some fine gravel and silts, and is usually well graded material, making it resistant to liquefaction process. The sand lenses are at a depth exceeding 60 feet and the effective overburden pressure is more than 5.5 kips per square foot with the addition of the plant foundation loads, the effective pressure on these lens could exceed by more than two times the pressure (4.1 kips per square foot), which according to Kishida (Reference 3) will prevent saturated sandy soils from liquefying. It is, therefore, concluded that sand lenses will not liquefy under the earthquake loading and consequently, no settlement due to liquefaction is anticipated.

CPS/USAR CHAPTER 02 B2.5-7 REV. 11, JANUARY 2005 B2.5.6 References

1. H. B. Seed and I. M. Idriss, "A Simplified Procedure for Evaluating Soil Liquefaction Potential," Report No. EERC 70-9, University of California, Berkeley, California, November 1970. 2. Shannon & Wilson, Inc. and Agbanian-Jacobsen Associates, "Soil Behavior Under Earthquake Loading Conditions: State of the Art Evaluation of Soil Characteristics for Seismic Response," prepared for the U.S. Atomic Energy Commission, January 1972. 3. H. Kishida, "Characteristics of Liquefied Sands During Mino-Owari, Tohnankai, and Fukui Earthquakes," Soils and Foundations (Japan), Vol. 9, No. 1, pp. 75-92, March

1969. 4. Benjamin and Cornell, "Probability, Statistics and Decision for Civil Engineers." 5. H. J. Gibbs and W. G. Holtz, "Research on Determining the Density of Sand by Spoon Penetration Testing," Proceeding of 4th International Conference on Soil Mechanics and Foundation Engineering, London, Vol. I, pp. 35-39, 1957. 6. E. D'Appolonia, "Dynamic Loadings," Journal of Soil Mechanics and Foundations Division, ASCE, No. SM1, Volume 96, p. 61, January 1970.

CPS/USAR CHAPTER 02 B2.5-8 REV. 11, JANUARY 2005 TABLE B2.5-1 SAND LENS SUMMARY ELEVATION (ft-in.) MATERIAL LENS NUMBER FROM TO UNIFIED CLASSIFICATION SOIL DESCRIPTION BORINGS CONSIDERED SPT REPRESENTATIVE BLOWCOUNT RELATIVE DENSITY** % FINER #200 SIEVE REMARKS 1 675 677.5 SP-SW Brown and gray P-32 9 (10 = D&M) 104.4*** N.A. 1) Tested and approved in med-coarse sand field with trace fine

2) Unit WT = 133.5 pcf dry sand and silt density 3) Confining pressure 2 677 679.5 SM Gray silty fine P-34 200 / 6 in. 95+ N.A. 1) Gradation to coarse sand

2) High R.D. with trace fine

3) Confining pressure gravel 3 652 665.5 SM-SP-SW Gray fine to P-10, P-14, P-30 68* 90+* 38-52% Based on statistical analysis coarse sand with P-33, P-34

, P-35 (From P-14, Will not liquefy some silt and P-36, P-37, P-41 P-15 1) Due to high R.D. fine gravel P-52, P-33A, P-15 Grain Size 2) Due to high % fines Analysis) 3) Due to gradation 4) Confining pressure 4 669 672.5 SW Gray fine to P-10 64 90 N.A. 1) High R.D. coarse sand with

2) Gradation trace silt

3) Confining pressure 5 663.5 760.5 SW-SM-SP Gray silty fine P-33, P-41, P-52 70 (AVG of 7) 95 1) High R.D. sand with some

2) Gradation fine to coarse

3) Confining pressure and fine gravel 6 661 667 SM Gray silty fine P-38, P-54 50 85 45 to 55 See Figure 2.5-17.29 sand 1) Will not liquefy due to high % fines 2) Confining pressure 3) R.D. _______________________________

• Based on statistical analysis with 97.5% probability.

• • Based on Gibbs & Holtz criteria.

• • • Based on inplace field testing.

CPS/USAR CHAPTER 02 C2.5-1 REV. 11, JANUARY 2005 ATTACHMENT C2.5 GEOLOGIC MAPPING PROGRAM FOR THE POWER BLOCK, ULTIMATE HEAT SINK, SCREEN HOUSE, AND SHUTDOWN SERVICE WATER SYSTEM EXCAVATIONS

CPS/USAR CHAPTER 02 C2.5-2 REV. 11, JANUARY 2005 ATTACHMENT C2.5 - GEOLOGIC MAPPING PROGRAM FOR THE POWER BLOCK, ULTIMATE HEAT SINK, SCREEN HOUSE, AND SHUTDOWN SERVICE WATER SYSTEM EXCAVATIONS C2.5.1 Introduction This attachment presents the results of a geologic mapping program conducted during the periods March 15-19, 1976, at the power block excavation; July 20-23, 1976, at the ultimate heat sink dam; May 24-26, 1977, at the baffle dike abutment; and June 16-23, 1977, at the screen house and outlet structure excavations. The purposes of the program were to verify the site stratigraphy as determined by the boring program and as described in the Preliminary Safety Analysis Report (PSAR), to gather data for preparation of the Final Safety Analysis Report (FSAR), and to confirm that the floor of the excavations extended into unaltered, Illinoian till as stated in the PSAR. Site stratigraphic units (all Pleistocene in age) discussed in this report are the Peyton Colluvium; the Cahokia Alluvium Henry Formation; the Richland Loess; the Wedron Formation; the Robein Silt; the weathered Glasford Formation; and the unaltered Glasford Formation. Descriptions and ages of each of the stratigraphic units are presented in Figure C2.5-1. The stratigraphic nomenclature has been refined from that presented in the PSAR: however, it does not indicate any difference between the lithologic units encountered in the borings and those mapped in the excavations. A comparison of terminology used in this Attachment and the PSAR is presented in Figure C2.5-2. Figure C2.5-3 shows locations of all major excavations referred to in the text. Figure C2.5-4 presents a plan view of the power block excavation showing the locations of main structures within the power block and the locations of the four excavation walls. Figure C2.5-5 shows the location of the screen house and the four excavation walls. Figure C2.5-6 gives a plan view of the ultimate heat sink area and shows the locations of the submerged dam, baffle dike, screen house, and shutdown

service water system (SSWS) outlet structure. C2.5.2 Conclusions The geologic mapping program confirmed that: a. The lithologic units exposed in the excavations are the same as those encountered in the borings and presented in the PSAR. One exception was the discovery of a black organic silt lens in the baffle dike abutment and submerged dam abutment excavations. b. The stratigraphic units are continuous across the excavations. c. Sand deposits within the tills, which were interpreted from the boring data as being discontinuous pockets or lenses, were exposed in the excavations and were confirmed to be discontinuous pockets. Two exceptions were noted: a nearly continuous, 2 to 3 feet thick, layer of brown, fine sand occurring within the Wedron Formation at approximately elevation 725 feet MSL in the power block and screen house excavations (Figures C2.5-7 through C2.6-12), and a continuous layer of fine to coarse sand at approximately elevation 658 feet MSL on the south abutment of the

submerged dam (Figure C2.5-13). d. The floor of the excavations extended into unaltered Illinoian till of the Glasford Formation as stated in the PSAR.

CPS/USAR CHAPTER 02 C2.5-3 REV. 11, JANUARY 2005 e. No geologic features showing vertical offset of stratigraphic units were noted. The foundation design assumptions presented in the PSAR are valid, and no changes based on soil conditions encountered in the excavations were required in the design of the power block and ultimate heat sink structures, with the exception of minor design changes at the abutments of the submerged dam and baffle dike.

C2.5.3 Scope The geologic mapping program consisted of establishing detailed descriptions of the site stratigraphic units; identifying, staking, and surveying contacts between these units at selected points within the excavations; and preparing geologic sections and photo mosaics (power block

excavation) to document the mapping program. T. C. Buschbach, H. B. Willman, D. L. Gross, L. R. Follmer, and C. S. Hunt of the Illinois State Geological Survey (ISGS) visited the site on Thursday, March 18, 1976, to inspect the stratigraphic units and contacts exposed in the power block excavation and to verify the results of the geologic mapping program. The ISGS agreed with Sargent & Lundy's identifications of the stratigraphic units and contacts exposed in the power block excavation (T. C. Buschbach, written communication to L.

J. Koch, Illinois Power Company, April 9, 1976). J. W. Skrove, U.S. Nuclear Regulatory Commission, visited the site on Tuesday, April 20, 1976, to inspect the stratigraphic units and contacts exposed in the power block excavation (J. W. Skrove, written communication to W. P. Gammill, DSE, April 27, 1976).

C2.5.3.1 Procedure The geologic mapping consisted mainly of a visual inspection of each of the excavation walls. Hand tools were used to trench into relatively undisturbed soil. In areas where localized sloughing had occurred near the floor of the power block excavation, a backhoe was used to dig trenches to expose the undisturbed soil. Descriptions of the stratigraphic units were prepared and stratigraphic contacts were staked and labelled. Geologic contacts separating the Wedron Formation, the Robein Silt, the weathered Glasford Formation, and the unaltered Glasford Formation were identified on the walls of the excavations. After identifying the stratigraphic units and contacts, the Sargent & Lundy mapping team located major sand deposits on the excavation walls. This was accomplished by visual observation of color and erosional patterns, investigation of areas of obvious groundwater seepage, and hand trenching on the excavation walls. After all staking had been completed, surveyors established the locations of the staked control points to the nearest 0.01 foot using plant grid coordinates and the elevations of the staked control points to the nearest 0.2 foot using mean sea level datum. C2.5.4 General Site Stratigraphy Stratigraphic units present at the power block, ultimate heat sink, screen house, and shutdown service water system (SSWS) excavations were the Peyton Colluvium; the Cahokia Alluvium Henry Formation; the Richland Loess; the Wedron Formation; the Robein Silt; the weathered Glasford Formation; and the unaltered Glasford Formation. Not a11 formations were present in all places.

See Section C.2.5.5 and Figures C2.5-1 and C2.5-2 for ages of the formations and lithologic details.

CPS/USAR CHAPTER 02 C2.5-4 REV. 11, JANUARY 2005 The Peyton Colluvium consists of brown clayey silt with minor amounts of gravel along the base of the slopes of the North Fork of Salt Creek. The colluvium was entirely removed from the submerged dam abutments, baffle dike abutment, screen house, and ECCS outlet structure excavations. The Cahokia Alluvium - Henry Formation consists of alluvial and outwash deposits and is confined to the valleys of Salt Creek and North Fork of Salt Creek. The alluvium is primarily poorly sorted silt, clay, and silty sand with lenses of sand and gravel. Underlying the alluvium is glacial outwash of the Henry Formation which consists of yellow-brown fine to coarse sand and gravel with pockets of gray and brown silt, sandy silt, and silty clay. A lag gravel is often present at the base of the Henry Formation. The contact between the Cahokia Alluvium and Henry Formation was not mapped because nearly all of the Cahokia Alluvium was removed from excavations where originally present and little of significance could be gained by mapping what remained. The contact between the Cahokia Alluvium - Henry Formation and Glasford Formation (weathered or unaltered) was identified by an abrupt change in color and decrease in grain size. The Richland Loess is almost entirely confined to the upland areas and consists of brown clayey silt, with trace fine sand. The loess had been removed or highly disturbed over most of the perimeter of the power block and screen house excavations prior to mapping. The contact with the underlying Wedron Formation was not identified. The Wedron Formation is almost entirely confined to the upland areas and consists of Wisconsinan till, which is brown to gray clayey silt to silty clay with some fine sand and trace gravel.

Discontinuous lenses of brown to gray, fine to coarse sand occur within the till. The contact between the Wedron Formation and the underlying Robein Silt was defined by an abrupt change in color, clay content, and presence of organic material. The Wedron Formation is a stable material on construction slopes of 1.5 horizontal to 1 vertical and, in general, makes excellent backfill. Locally, the glacial till includes sand pockets and lenses generally with water seeps. The sands are easily eroded by water flowing down the slope. The Robein Silt consists of water-deposited loess characteristically with dark brown organic silt and traces of clay and fine sand. Locally, some peat is also found. The Robein Silt is confined to the upland areas. The Robein Silt has a tendency to break into vertical slopes, especially under freeze-thaw conditions. Included organic material and clay, acting as a binder, reduces erosion by running water

to a minimum. The weathered Glasford Formation consists of weathered Illinoian till that was weathered during Sangamonian and later times. It is characterized by gray silt, with trace fine sand grading to gray-green silty clay or clayey silt, with some fine to coarse sand and trace fine to coarse gravel. The weathered Glasford Formation is absent in the valleys. The weathered till is slightly to highly calcareous. The contact within the Glasford Formation, between the weathered and unaltered till, is gradational. Based on the site visit by the ISGS, this contact was defined where a very strong reaction to dilute muriatic acid (HCl) was achieved, where no noticeable color change occurred with increased depth, and where a normal gravel fraction was present in the till. The weathered Glasford Formation is stable on slopes of 1.5 horizontal to 1 vertical. A few seeps, associated with discontinuous sand lenses, were noted.

CPS/USAR CHAPTER 02 C2.5-5 REV. 11, JANUARY 2005 The unaltered Glasford Formation consists of unaltered Illinoian till and is characterized by a dark gray clayey silt with some interspersed fine to coarse sand and fine to coarse gravel. The till is highly calcareous and contains discontinuous lenses of gray, fine to coarse sand. Unweathered Glasford Formation is present everywhere on the site. Unweathered Glasford Formation is a very hard material stable on slopes of 1.5 horizontal to 1 vertical. Some sand pockets were noted. In the ultimate heat sink excavation, a silt layer similar to the Robein Silt was exposed at about elevation 660, some 20 feet below the top of the unweathered Glasford Formation.

C2.5.5 Findings C2.5.5.1 Power Block Excavation Stratigraphic units presented at the power block excavation, in descending order, were: the Richland Loess, highly disturbed due to construction operations; the Wedron Formation; the Robein Silt; the weathered Glasford Formation; and the unaltered Glasford Formation. Figures C2.5-7 through C2.5-10 show that, with the exception of the southwest corner of the excavation, the contact between the Wedron Formation and the underlying Robein Silt was generally continuous at approximately elevation 698 feet MSL. In the southwest corner (Figures C2.5-9 and C2.5-10), the Robein Silt and the underlying, weathered Glasford Formation had been eroded almost completely away by stream action during the early part of the Wisconsinan Age. The contact between the weathered and unaltered portion of the Glasford Formation was at a fairly constant elevation of approximately 685 feet MSL. Sand deposits were exposed in the walls of the excavation within the Wedron Formation and near the contact between the weathered and the unaltered portions of the Glasford Formation. The sand deposits within the Wedron Formation consisted of a nearly continuous seam of brown, fine sand generally 2 to 3 feet thick that spanned all four walls of the excavation at approximately elevation 725 feet MSL and thin, discontinuous seams and localized pockets of gray, fine to coarse sand within the till, extending several feet vertically and horizontally. Sand deposits were difficult to identify on the south and east excavation walls due to sloughing and continuing excavation by dragline. The geologic sections in Figures C2.5-7 through C2.5-10 show both the sand deposits identified in the field and those established from the interpretation of photographs. Sand deposits were exposed near the contact between the weathered and the unaltered portion of the Glasford Formation on the east, south, and west walls of the excavation. These sand deposits were variable in texture and discontinuous in extent and carried varying amounts of groundwater into the excavation. Along the east wall (Figure C2.5-8), reference points marked with "X's" within the unaltered Glasford Formation at stations 4+71 North and 7+49 North mark the locations of small sand lenses from which groundwater was seeping at a perceivable rate into the excavation. The floor of the excavation was located entirely within the unaltered Illinoian till of the Glasford Formation in all areas inspected by the mapping team. The elevation of the floor varied across the excavation because of over-excavation to remove sand pockets within the unaltered Glasford Formation and to provide drainage to a common sump. The entire floor of the excavation had not been cleaned at the time of field mapping to provide a fresh exposure of the unaltered Illinoian till of the Glasford Formation. A systematic testing program and visual inspection by onsite personnel were performed to verify the presence of the unaltered Glasford Formation on the floor of the excavation prior to fill placement.

CPS/USAR CHAPTER 02 C2.5-6 REV. 11, JANUARY 2005 C2.5.5.2 Screen House Excavation Stratigraphic units exposed on the walls of the screen house excavation included the Cahokia Alluvium - Henry Formation; the Wedron Formation; the Robein Silt; and the Glasford Formation, consisting of a weathered and unaltered zone (see Figures C2.5-11, C2.5-12, C2.5-14, and C2.5-15). Richland Loess (not mapped) was exposed locally, but was highly disturbed by construction activities. The Cahokia Alluvium - Henry Formation is exposed on the south and west walls (Figures C2.5-14 and C2.5-15) of the screen house excavation, where the excavation extends into the flood plain of the North Fork of Salt Creek. Mapping was performed after the ultimate heat sink was excavated to elevation 668.5 feet MSL; consequently, the upper part of the Cahokia Alluvium - Henry Formation was not exposed at the time of mapping. The alluvial and glacial outwash deposits occur from the surface at elevation 680 feet MSL to approximately elevation 660 feet MSL, where they are underlain by the unaltered Glasford Formation. On the north and east walls of the screen house excavation (Figures C2.5-11 and C2.5-12), the youngest stratigraphic unit was the Wedron Formation. In the Wedron Formation, a nearly continuous seam of brown, fine sand, generally 1 to 2 feet thick was exposed on the east wall of the excavation at approximately 725 feet MSL. Thin, discontinuous seams and lenses of gray, fine to coarse sand were noted within the Wedron Formation. The contact between the Wedron Formation and the underlying Robein Silt was generally continuous at approximately elevation 686 feet MSL. Illinoian till of the Glasford Formation, consisting of a weathered and unaltered zone, underlies the Robein Silt at approximately elevation 684 feet MSL. The contact between the weathered and unaltered zones was gradational at approximately elevation 680 feet MSL. Discontinuous sand units were exposed near the contact between the weathered and unaltered Glasford Formation on the east and north walls of the excavation. Groundwater flow from these sand units was minor and variable. The floor of the excavation at elevation 653 feet MSL was located entirely within the unaltered till. Some overexcavation was necessary to remove discontinuous sand lenses and to provide drainage to a common sump. C2.5.5.3 Ultimate Heat Sink C2.5.5.3.1 Submerged Dam and Baffle Dike Excavations Excavations for the submerged dam and baffle dike in the heat sink bottom revealed the Cahokia Alluvium - Henry Formation overlying hard glacial till of the unaltered Glasford Formation. Mapping was performed after the heat sink bottom was excavated to elevation 668.5 feet MSL; consequently, the upper part of the Cahokia Alluvium - Henry Formation was not present at the time of mapping.

Geologic profiles (Figures C2.5-16 and C2.5-17) indicate the contact between the Cahokia Alluvium - Henry Formation and the underlying unaltered Glasford Formation is continuous at approximately elevation 658 feet MSL for the submerged dam and approximately elevation 660 feet MSL for the baffle dike. C2.5.5.3.2 Submerged Dam Abutments Stratigraphic units exposed in the abutments of the submerged dam in descending order are: the Richland Loess, mappable only on the south abutment; the Wedron Formation; the Robein Silt; and the Glasford Formation, consisting of weathered and unaltered zones. Figures C2.5-18 and C2.5-19 present a plan view of the north abutment and south abutment, respectively.

CPS/USAR CHAPTER 02 C2.5-7 REV. 11, JANUARY 2005 The contact between the Richland Loess and the Wedron Formation varies with the topography along the south abutment (Figure C2.5-13). The contact between the Wedron Formation and the underlying Robein Silt is continuous at approximately elevation 696 feet MSL for the south abutment and elevation 702 feet MSL for the north abutment (Figure C2.5-20). The contact between the Robein Silt and the underlying weathered portion of the Glasford Formation is generally continuous at approximately elevation 694 feet MSL for the south abutment and elevation 699 feet MSL for the north abutment. The contact between the weathered and unaltered portions of the Glasford Formation varies due to the proximity to the natural ground surface. Geologic mapping verified unaltered Glasford Formation below elevation 676 feet MSL. Sand deposits are exposed in the abutments within the Wedron and Glasford Formations. The sand deposits within the Wedron Formation consist of local pockets of brown, fine to coarse sand. The sand deposits within the Glasford Formation are generally discontinuous seams and localized pockets of gray, fine to coarse sand. Between elevation 660 and 664 feet MSL, silt and sand seams, interbedded with a 1 to 2 foot thick layer of organic silt, were noted within the unaltered Glasford Formation. Seepage of groundwater was noted within the sand directly above the

unaltered Glasford Formation. On the north abutment the sand seams were not present and organic silt at approximately elevation 663 MSL was overlain by unaltered Glasford Formation and was underlain by a thin strata of weathered Glasford Formation, which grades into unaltered Glasford Formation. C2.5.5.3.3 Baffle Dike Abutment Stratigraphic units exposed in the baffle dike abutment in descending order were: the Richland Loess, which was highly disturbed due to construction operations; the Wedron Formation; the Robein Silt; and the Glasford Formation, consisting of weathered and unaltered zones. Figure C2.5-21 shows a geologic section of the baffle dike abutment. The contact between the Wedron Formation and the underlying Robein Silt is continuous at approximately elevation 687 feet MSL. The Robein Silt is underlain by the weathered Glasford Formation at an elevation of about 686 feet MSL. The contact between the weathered and unaltered zones of the Glasford Formation was

mapped at approximately elevation 676 feet MSL. The unaltered Glasford Formation between elevation 676 feet MSL and 668 feet MSL was soft to medium stiff, and the till below elevation 668 feet MSL was hard. A large sand lens in the Glasford Formation pinches out on the north side of the baffle dike abutment but continues on the south side of the dike. Groundwater emerged from the lower part of this sand deposit. Two thin discontinuous lenses of black organic silt were noted at approximately elevations 662 feet MSL (upper) and 651 feet MSL (lower) near the large sand lens at the north and south ends of the excavation. C2.5.5.4 SSWS Outlet Structure Stratigraphic units exposed to the SSWS outlet structure excavation include: Peyton Colluvium, which is exposed at the base of the bluffs of the North Fork of Salt Creek; the Cahokia Alluvium - Henry Formation; and the Glasford Formation. Figure C2.5-22 is a geologic section of the east wall of the excavation; the other walls of the excavation revealed Cahokia Alluvium - Henry Formation deposits underlain by unaltered Glasford Formation. Peyton Colluvium was mapped along a sloping contact with the underlying Glasford Formation between elevation 672 and 679 feet MSL. Within the Glasford Formation at the SSWS outlet structure excavation were two organic silt units. The upper silt unit was exposed near the contact CPS/USAR CHAPTER 02 C2.5-8 REV. 11, JANUARY 2005 with the Peyton Colluvium from elevation 670 feet MSL (north) to elevation 678 feet MSL (south). The lower silt unit occurred at approximately elevation 662 feet MSL, near the floor of the excavation. C2.5.5.5 SSWS Piping Subgrade The SSWS piping excavations extended from the screen house and SSWS outlet structure to the south side of the power block excavation (Figure C2.5-23). The subgrade for the SSWS pipes consists mainly of Wisconsinan till of the Wedron Formation. Near the screen house and SSWS outlet structure, the Robein Silt was exposed in the piping excavation at approximately elevation 688 feet MSL. The silt was partially removed by overexcavation and the excavation was raised to the proper level with suitable backfill material. Several thin, discontinuous sand and silt lenses and small randomly distributed sand and silt pockets

were noted in the Wedron Formation.

CPS/USAR CHAPTER 02 D2.5-1 REV. 12, JANUARY 2007 ATTACHMENT D2.5 UNPUBLISHED NOTES

CPS/USAR CHAPTER 02 D2.5-2 REV. 11, JANUARY 2005 PITTSFIELD-HADLEY ANTICLINE The Pittsfield-Hadley Anticline is a prominent structure that crosses the Mississippi River Arch in a northwest-southeast direction. It is located in Lewis County, Missouri, and Pike County, Illinois. It plunges southeastward and it appears to lose its identity in Greene County, Illinois. Pennsylvanian strata on the flanks at the anticline dip somewhat less than underlying Mississippian strata. This indicates that some uplift occurred during post-Mississippian and pre-Pennsylvanian time, but the major uplift took place after Pennsylvanian time. Total uplift along the anticline exceeds 300 feet in some places. Folds with similar directional trends, but with uplift of only slightly more than 100 feet occur at Fishhook in Pike County and at Media in Henderson County, both in Illinois. In view of the similarities of their orientation and their stratigraphy, these minor structural highs are assumed to have formed at the same time as did the Pittsfield-Hadley anticline.

T. C. Buschbach 5/73 Ref. F. Krey, Structural Reconnaissance of the Mississippi Valley Area from Old Monroe. Missouri to Nauvoo, Illinois: Illinois Geol. Survey Bull. 45., 1924. MISSISSIPPI RIVER ARCH The Mississippi River Arch is a broad, corrugated fold which extends generally north-south through the bulge of western Illinois. To the north it blends with the Wisconsin uplands and to the south it intercepts the Lincoln Anticline. The arch separates the Illinois Basin from the Forest City Basin.

Dating of movements along the arch is difficult because erosion has removed the Pennsylvanian strata. However, it appears that the Mississippi River Arch existed early in Pennsylvanian time and was probably subjected to additional deformation at the end of Paleozoic time. The arch is cut by numerous cross folds which trend northwest-southeast and plunge gently southeastward in the Illinois Basin.

T. C. Buschbach 5/73 Ref. J. V. Howell, The Mississippi River Arch: Kans. Geol. Soc. Guidebook, 9th Annual Field Conf., pp. 386-389, 1935. SALEM-LOUDEN ANTICLINAL BELT The Salem-Louden Anticlinal Belt is a prominent structural high in the Illinois Basin. The anticlinal belt trends northnortheast and is most prominent in Central Marion and eastern Fayette Counties. Closure of 100 feet or slightly more is common on the anticline, and the structure has proved to be important in determining the position of oil fields. Individual units within the Pennsylvanian System thin over the top of the anticlinal belt, indicating that the structure was uplifted during Pennsylvanian time. Uplift also continued after Pennsylvanian deposition ended.

T. C. Buschbach 5/73 Ref. E. P. DuBois, Geology and Coal Resources of a Part of the Pennsylvanian System in Shelby, Moultrie, and Portions of Effingham and Fayette Counties: Illinois Geol. Survey Rept. Inv. 156, 1951.

CPS/USAR CHAPTER 02 D2.5-3 REV. 11, JANUARY 2005 SANGAMON ARCH The Sangamon Arch was formed by uplift in central and western Illinois during Devonian and early Mississippian time. The arch extends from the Mississippi River Arch eastward to Macon and DeWitt Counties in central Illinois. Its limits are reasonably well defined by the zero isopach of the Cedar Valley (Middle Devonian) Limestone. Although several hundreds of feet of Devonian and Silurian strata normally present in surrounding areas were either not deposited over, or eroded from the arch, later movements have masked the arch so that it does not show on structure maps of the area. It is a relict structure that is interpreted by the stratigraphy of the region.

T. C. Buschbach 5/73 Ref. L. L. Whiting, and D. L. Stevenson, The Sangamon Arch: Illinois Geol. Survey Circ. 383, 20 pp., 1965. DUPO-WATERLOO ANTICLINE The Dupo-Waterloo Anticline strikes north-northwest from Monroe County, Illinois, through St. Louis, Missouri, and appears to terminate against the Cap au Gres faulted flexure about 12 miles north of St. Louis. Outcrops in the Dupo area show that the east flank dips 2 to 3 degrees, whereas the west flank dips up to 30 or more degrees. The Dupo-Waterloo anticline was probably intermittently active from Silurian time to post-Pennsylvanian time. Major movement appears to have occurred near late Mississippian, pre-Pennsylvanian time, with renewed uplift in post-Pennsylvanian, pre-Pleistocene time. Total strucutral relief is at least 500 feet near Waterloo. (Precambrian high?)

T. C. Buschbach 5/73 Ref. A. H. Bell, The Dupo Oil-Field: Illinois Geol. Survey, Illinois Petroleum 17, 1929.

SANDWICH FAULT ZONE The Sandwich Fault Zone strikes across northern Illinois in a west-northwest direction from western Will County to Ogle County. It forms the northern boundary of the Ashton Arch. The fault zone has a maximum downthrow of at least 900 feet on the northeast side. The throw decreases toward its eastern end, and a scissors effect causes the southwest end of the fault to be downthrown a little more than a hundred feet in western Will County. Movements along the Sandwich Fault Zone were post-Silurian, pre-Pleistocene. No rocks representing the intervening time are present in the area. However, major movements along the fault zone may have occurred at about the same times that the La Salle Anticlinal Belt was uplifted-in post-Mississippian, pre-Pennsylvanian time and again in

post-Pennsylvanian times.

T.C. Buschbach 5/73 Ref. H. B. Willman, and J. N. Payne, Geology and Mineral Resources of the Marseilles, Ottawa, and Streator Quandrangles: Illinois Geol. Survey Bull. 66, 1942. H. B. Willman, and J. S. Templeton, Cambrian and Lower Ordovician Exposures In Northern Illinois: Trans. of Illinois Acad. Sci. v. 44, p. 109-125, 1951.

CPS/USAR CHAPTER 02 D2.5-4 REV. 11, JANUARY 2005 LASALLE ANTICLINAL BELT The La Salle Anticlinal Belt is an extensive asymmetrical fold that extends in Illinois from Lee County in the northwest to Lawrence County in the southeast. The west limb dips sharply into the deeper part of the Illinois Basin, whereas the east limb dips gently into the eastern shelf area of the basin. The crest of the anticline plunges to the south-southeast. Initial deformation along the La Salle Anticlinal Belt took place in post-Mississippian time. Deformation continued through early Pennsylvanian time, particularly at the southern part of the structure. Renewed activity occurred after Pennsylvanian time, probably at the close of the Paleozoic Era.

T. C. Buschbach 5/73 Refs. G. H. Cady, The Structure of the La Salle Anticline: Illinois Geol. Survey Bull. 36, p. 171-177, 1920.

J. N. Payne, The Age of the La Salle Anticline: Trans Illinois Acad. Sci. Vol. 32, No. 2, p. 5-7, 1939.

CENTRALIA FAULT The Centralia Fault strikes nearly north-south parallel to, and one mile east of the DuQuoin Monocline in Marion and Jefferson Counties. It is a zone of several parallel faults. Net displacement is downward to the west, with maximum displacement of about 200 feet. The faults can be seen in several coal mines in the Centralia area, but they are not visible at the land surface. The faults appear to be the results of shearing stresses formed after folding took place on the DuQuoin Monocline. Relief of the stresses was upward on the east side, opposed to the east dip of the monocline. The faulting occurred in post-Pennsylvanian, pre-Pleistocene time.

T. C. Buschbach 5/73 Ref. R. L. Brownfield, Structural History of the Centralia Area: Illinois Geol. Survey Rept. Inv. 172, 1954. ROUGH CREEK LlNEAMENT The Rough Creek Lineament is a series of faults and fault zones extending generally east-west through western Kentucky and southern Illinois. In Kentucky, it includes the Rough Creek Fault Zone. In Illinois, it includes the east-west portion of the Shawneetown Fault Zone to the east and the Cottage Grove Fault System to the west. Heyl (1972, p. 885) suggests that strike-slip faulting or wrench faulting is a major component in the Rough Creek Lineament. He tentatively includes it in a line or zone of faults, monoclines, and igneous intrusions. The line extends east-west for 800 miles along the 38th parallel from West Virginia to at least as far west as the Ozark Uplift. In Illinois, the lineament includes numerous high angle reverse faults with the south side upthrown. They appear to be the result of compressional forces from the south, and they display evidence of some horizontal movements. The eastern part of the lineament, the Shawneetown Fault Zone, is dominated by high angle thrust faulting. Displacement is locally as great as 3400 feet and may be considerably more. The Shawneetown extends westward along the prominent hills in southern Gallatin County, curves southward around Cave Hill in Saline County, leaves the Rough Creek Lineament and joins the southwest-trending Herod Fault to form the Lusk Creek Fault Zone. The western portion of the lineament, the Cottage Grove Fault System, appears to have formed at roughly the same time as the Shawneetown, but displacements are much diminished, with maximum CPS/USAR CHAPTER 02 D2.5-5 REV. 11, JANUARY 2005 ROUGH CREEK LlNEAMENT (Continued) displacements of about 250 feet. From all available evidence it appears that the age of faulting along the Rough Creek Lineament is chiefly post-Pennsylvanian, pre-late Cretaceous, although some workers have suggested the possibility of later movements because of recent seismic activity in the general area.

T. C. Buschbach 5/73 Ref. A. V. Heyl, The 38th Parallel Linement and Its Relationship to Ore Deposits: Economic Geology, Vol. 67, No. 7 pp. 879-894. 1972. WABASH VALLEY FAULT SYSTEM The Wabash Valley Fault System is a series of generally parallel faults that trends northeastward from the Rough Creek Lineament in Gallatin County, roughly paralleling the Wabash River in Illinois and Indiana to near Mt. Carmel, Wabash County, Illinois. The faults are high angle, normal faults.

The faults have been observed in mines, boreholes, and surface exposures. Maximum known displacement on the faults is a little over 400 feet, although displacement of a few to 200 feet are more common. The throw of the faults appears to be the same in Mississipian and Pennsylvanian strata,and they are clearly post-Pennsylv anian in age. No displacement has been recognized in Pleistocene deposits, thus the faulting appears to have occurred in pre-Pleistocene time.

T. C. Buschbach 5/73 Refs. J. A. Harrison, Subsurface Geology and Coal Resources of the Pennsylvania System in White County, Illinois: Illinois Geol. Survey Rept. Inv. 153, 1951. D. H. Swann, Waltersburg Sandstone Oil Pools of Lower Wabash Area, Illinois and Indiana: Bull. Amer. Assoc. Pet. Geologists, Vol. 35, No. 12, 1951. STE. GENEVIEVE FAULT SYSTEM The Ste. Genevieve Fault System extends northwestward across Union County, Illinois, crossing the Mississippi River just north of Grand Tower. It continues in that direction through Perry County, Missouri, and then swings westward through Ste. Genevieve County. Although the system includes numerous horsts and grabens, the net displacement is down to the north and east. Maximum displacement is more than 1000 feet and may approach 2000 feet. The fault system forms a sharp boundary, a few miles wide, between the Illinois Basin and the Ozark Uplift. The faults are high angle faults with some reverse and some lateral movements. Compression from the southwest was probably an important factor in their formation. Movements along the faults occurred several times during late Paleozoic times. Without question, there was post-Mississippian, pre-Pennsylvanian movement followed by post-Pennsylvania movement. Unusually thick sections of Devonian strata are preserved in grabens of the fault system, and these may be explained by earlier faulting in perhaps Late Devonian time.

The extension of the Ste. Genevieve Fault System into Illinois has been called the Rattlesnake Ferry Fault. Presently it is called the Ste. Genevieve Fault Zone.

T. C. Buschbach 5/73 CPS/USAR CHAPTER 02 D2.5-6 REV. 11, JANUARY 2005 STE. GENEVIEVE FAULT SYSTEM (Continued)

Refs. S. St. Clair and S. Weller, Geology of Ste. Genevieve County, Missouri: Missouri Bur. Geology and Mines, Ser. 2, V. 22, 1928. W. F. Meents and D. F. Swann, Grand Tower Limestone (Devonian) of Southern Illinois: Illinois Geol. Survey Circ. 389, 1965. CAP AU GRES FAULTED FLEXURE The Cap Au Gres flexure is a sharp monoclinal fold that extends east-south-eastward through Lincoln County, Missouri, then generally eastward through southern Calhoun and Jersey Counties, Illinois. The rocks dip steeply on the southern flank of the structure, and the maximum amount of structural relief is 1000 to 1200 feet. Along much of the length of the flexure the rocks are broken by faults that trend parallel to the strike of the rocks. The faults generally are down thrown to the south and have displacements from a few to a few hundred feet. Limited exposures in the area make it difficult to determine the extent and continuity of the faults. Major deformation along the Cap Au Gres Flexure took place in post-Middle Mississippian, pre-Pennsylvanian time. Pennsylvanian strata south of the flexure are considerably lower than outliers of similar strata north of the Flexure. In addition, the Calhoun peneplain bevels across the edges of tilted Pennsylvanian strata in the area, thus indicating post-Pennsylvanian movement along the Cap Au Gres Faulted Flexure. Ruby (1952, pp. 64, 145, 146) argues that even later movement is indicated by displacement along the structure of the Calhoun peneplain and the Grover Gravel, which immediately overlies the peneplain. This displacement occurred after deposition of the gravel (probably in Pliocene time) and amounts to a little more than 100 feet. No evidence has been found to indicate any deformation of Pleistocene deposits in the area. A pair of northwest-trending anticlines end abruptly against the flexure; they are the Dupo-Waterloo Anticline to the south and the Lincoln Fold to the North. Both anticlines have their steeper flanks to the west, and they appear to have similar geologic histories. The crests of the anticlines are offset about 30 miles. It is possible that the Cap Au Gres Faulted Flexure is a left-lateral fault in which the horizontal movement of about 30 miles offset the Lincoln Fold from its southern continuation, the Dupo-Waterloo Anticline.

T. C. Buschbach 5/73 Revised 9/75 Ref. W. W. Ruby, Geology and Mineral Resources of the Hardin and Brussels Quadrangles (in Illinois): U.S. Geol. Survey Prof. Paper 218, 1952.