Text

Redacted St. Lucie, Unit 1, Amendment 29 to Updated Final Safety Analysis Report, Chapter 2, Site Characteristics
	ML18320A264
Person / Time
Site:	Saint Lucie
Issue date:	10/04/2018
From:	; Florida Power & Light Co
To:	; Office of Nuclear Reactor Regulation
References
	L-2018-172
	Download: ML18320A264 (952)
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TABLE 2.3-16 8 HOUR AT 5 miles FLORIDA POWER & LIGHT COMPANY CODE*LSD-4 St. Lucie Plant Site, Florida HIGHEST CALCULATED AVERAGE RELATIVE CONCENTRATION BY WIND DIRECTION AND CODED DATE FOR THE TIME PERIOD OF 8 HOURS AT THE LOW POPULATION DISTANCE OF 5 MILES PERIOD OF RECORD: 03/01/71 TO 02/29/72

LAST HOUR WIND CODED DATE MAXIMUM DIRECTION MONTH - DAY - HOUR X/Q (SEC/M**3)

NNE 2 29 02 2.021D-06

NE 3 22 13 2.292D-06

ENE 6 26 08 5.793D-06

E 5 18 09 3.518D-06

ESE 5 05 09 7.969D-06

SE 6 29 04 3.001D-06

SSE 8 18 05 3.241D-06

S 12 31 08 3.010D-06

SSW 3 10 06 4.105D-06

SW 11 27 24 4.280D-06

WSW 3 12 04 2.766D-06

W 11 20 09 2.284D-06

WNW 2 26 09 3.965D-06

NW 2 01 08 3.344D-06

NNW 3 09 03 3.947D-06

N 2 29 09 2.874D-06

2.3-47 TABLE 2.3-17 16 HOUR AT 5 miles FLORIDA POWER & LIGHT COMPANY CODE*LSD-4 St. Lucie Plant Site, Florida HIGHEST CALCULATED AVERAGE RELATIVE CONCENTRATION BY WIND DIRECTION AND CODED DATE FOR THE TIME PERIOD OF 16 HOURS AT THE LOW POPULATION DISTANCE OF 5 MILES PERIOD OF RECORD: 03/01/71 TO 02/29/72

LAST HOUR WIND CODED DATE MAXIMUM DIRECTION MONTH - DAY - HOUR X/Q (SEC/M**3)

NNE 2 29 10 1.074D-06

NE 3 22 13 1.146D-06

ENE 6 26 11 3.059D-06

E 5 18 09 2.132D-06

ESE 5 05 15 4.262D-06

SE 6 29 18 1.519D-06

SSE 8 18 06 1.713D-06

S 12 31 10 1.561D-06

SSW 3 10 14 2.056D-06

SW 11 28 09 2.275D-06

WSW 6 11 09 1.413D-06

W 11 20 12 1.146D-06

WNW 2 26 09 1.982D-06

NW 10 12 07 2.460D-06

NNW 3 09 10 2.484D-06

N 2 29 09 1.437D-06

2.3-48 TABLE 2.3-18 72 HOUR AT 5 miles FLORIDA POWER & LIGHT COMPANY CODE*LSD-4 St. Lucie Plant Site, Florida HIGHEST CALCULATED AVERAGE RELATIVE CONCENTRATION BY WIND DIRECTION AND CODED DATE FOR THE TIME PERIOD OF 72 HOURS AT THE LOW POPULATION DISTANCE OF 5 MILES PERIOD OF RECORD: 03/01/71 TO 02/29/72

LAST HOUR WIND CODED DATE MAXIMUM DIRECTION MONTH - DAY - HOUR X/Q (SEC/M**3)

NNE 9 30 03 3.030D-07

NE 4 13 05 4.529D-07

ENE 6 27 09 6.837D-07

E 5 19 24 6.781D-07

ESE 5 05 22 9.630D-07

SE 7 01 21 6.787D-07

SSE 7 26 01 6.219D-07

S 1 12 11 7.025D-07

SSW 3 11 04 5.963D-07

SW 6 25 14 7.514D-07

WSW 1 31 02 5.992D-07

W 6 12 04 4.679D-07

WNW 10 23 07 6.140D-07

NW 2 23 07 6.840D-07

NNW 3 09 22 6.751D-07

N 2 29 09 4.348D-07

2.3-49 TABLE 2.3-19 26 DAY AT 5 miles FLORIDA POWER & LIGHT COMPANY CODE*LSD-4 St. Lucie Plant Site, Florida HIGHEST CALCULATED AVERAGE RELATIVE CONCENTRATION BY WIND DIRECTION AND CODED DATE FOR THE TIME PERIOD OF 624 HOURS AT THE LOW POPULATION DISTANCE OF 5 MILES PERIOD OF RECORD: 03/01/71 TO 02/29/72

LAST HOUR WIND CODED DATE MAXIMUM DIRECTION MONTH - DAY - HOUR X/Q (SEC/M**3)

NNE 7 29 02 1.045D-07

NE 8 03 08 1.278D-07

ENE 6 26 11 1.753D-07

E 6 09 08 2.259D-07

ESE 5 28 08 2.461D-07

SE 7 24 23 2.356D-07

SSE 8 06 05 2.522D-07

S 1 25 11 3.682D-07

SSW 3 26 24 1.853D-07

SW 7 18 08 2.434D-07

WSW 6 30 03 1.849D-07

W 6 22 17 1.636D-07

WNW 2 26 09 1.931D-07

NW 2 26 09 2.548D-07

NNW 2 28 01 2.022D-07

N 2 29 09 1.443D-07

2.3-50

TABLE 2.3-93 ANNUAL ST. LUCIE RELATIVE CONCENTRATION VALUES FOR SELECTED WORST PERCENTAGES

Accident Criteria Distance 1971* 1972 1973 0-2 hours worst 5% 1555 meters 8.55 D-05** 1.19 D-04 8.15 x D-05 0-2 hours worst 50% " " 8.58 D-07 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months worst 8045 meters 7.97 D-06 7.47 D-06 9.97 x D-06 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months worst 5% " " 4.83 D-07 4.96 D-07 5.59 x D-07 16 hours1.851852e-4 days 0.00444 hours 2.645503e-5 weeks 6.088e-6 months worst " " 4.26 D-06 3.73 D-06 5.86 x D-06 16 hours1.851852e-4 days 0.00444 hours 2.645503e-5 weeks 6.088e-6 months worst 5% " " 4.48 D-07 4.68 D-07 4.28 x D-07 3 days worst " " 9.63 D-07 1.06 D-06 1.70 x D-06 3 days worst 5% " " 3.14 D-07 3.42 D-07 2.87 x D-07 26 days worst " " 3.68 D-07 3.29 D-07 4.08 x D-07 26 days worst 5% " " 1.96 D-07 1.96 D-07 2.02 x D-07

1971 Period of Record is 3/1/71 to 2/29/72

- D-05 denotes x 10

-5

2.3-78 DELETED FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT UNIT 1 FIGURE 2.3-1 Amendment No. 23 (11/08)

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DELETED FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT UNIT 1 FIGURE 2.4-1 Amendment No. 23 (11/08)

Withheld Under 10 CFR 2.390

2.5 GEOLOGY

AND SEISMOLOGY

2.5.1 BASIC

GEOLOGIC AND SEISMIC INFORMATION

The St. Lucie site is located on Hutchinson Island on the east coast of

peninsular Florida, approximately 10 miles southeast of Fort Pierce, as

shown on Figure 2.5-1.

Subsurface investigations were performed at the site during 1967 and 1968

in conjunction with the PSAR studies for St. Lucie Unit 1. The activities

performed at the time included: literature studies, surface mapping, aerial

photographic studies, geophysical studies, and 78 borings. In 1973 the

Unit 1 FSAR and Unit 2 PSAR were submitted. Following AEC review in 1973, questions were posed concerning geology and seismology. Further

investigations were performed during 1973 and 1974, including: air photo

studies, additional literature review, 7 borings, and 50 miles of

continuous seismic reflection survey (See Appendix 2E). This current

report on the geology and seismology of the St. Lucie site incorporates all

data obtained in the investigations and is presented as specified in the, "Standard Format and Content of Safety Analysis Reports for Nuclear Power

Plants," as revised February, 1974.

The original surficial deposits at the plant site were excavated to

elevation -60 feet and backfilled with Category I fill. The Category I

structures will be supported on mat foundations within this fill.

The fill is underlain by the Anastasia Formation, a sequence of partially

cemented sands and sandy limestones, which extend to an average elevation

of about -145 feet. The Anastasia is underlain to an elevation of about

-600 feet to -700 feet by the partially cemented and indurated sands, clays

and sandy limestones of the Hawthorne Formation. Underlying these surface

strata are about 13,000 feet of Jurassic through Tertiary Formations, primarily carbonate rocks. These formations have a relatively gentle slope

to the southeast.

2.5.1.1 Regional Geology

The following discussion of regional geology is based on a comprehensive

review of available data, including published and unpublished reports and

maps, and interviews with recognized authorities. The study region

includes peninsular Florida with major emphasis placed on the subregional

area, as shown on Figure 2.5-1.

2.5.1.1.1 Physiography

The Floridan Plateau is a partially submerged peninsula of the North

American Continental Shelf. The peninsula of the State of Florida is the

exposed portion above sea level of the Floridan Plateau, and lies entirely

within the Coastal Plain physiographic province.

The major land forms are recognized from studies of aerial photographs and

are evident on topographic maps. They are generally aligned in a northerly

direction and may be grouped into three classifications: 1) highland

2.5-1 Amendment No. 25 (04/12) ridges, 2) interior plains and valleys, and 3) coastal lowland areas. The extent of these general land forms are shown on Figure 2.5-2.

The highland ridges are erosional remnants of Pliocene age deltaic and

terrestrial stream deposits. These remnants are the most prominent

topographic features in Florida. They occur as north-south aligned ridges

located in central and western peninsular Florida with surface elevations

generally ranging from 150 to 200 feet. These ridges may represent

erosional remnants of a former broad plain (White 1970). Most of these

ridges show evidence of advanced solution activity in the limestones below.

The remaining portion of peninsular Florida is mainly covered by

Pleistocene age terraces or contemporaneous deposits. The portion of

Florida covered by Pleistocene deposits can be subdivided into the interior

plains and valleys and the coastal lowlands. The surface elevation of the

interior plains and valleys ranges from about 40 to 150 feet, with the

highest areas occurring adjacent to the erosional remnants of the highland

ridges. This area is characterized by broad terraced plains and some

solutioned interior valleys.

The coastal lowlands are characterized by relatively flat relief and swampy

or marsh terrain. The lowlands along the east coast are low marine

terraces at an average elevation of about 25 feet. The lowlands in south

Florida are broad swamps and marshes. The lowlands along the west coast (generally north of Tampa) are characterized by relatively thin sand

deposits overlying Tertiary limestones. This region displays numerous

shallow depressions as a result of the solutioning of the relatively

shallow underlying limestones. Along the east coast, the surface sands are

thicker and shallow surface depressions are not as evident.

2.5.1.1.2 Geologic History

Paleozoic rocks underlie all of Florida, forming the basement complex on

which younger formations were deposited. However, due to the lack of

exposure of these older strata, hypotheses regarding Paleozoic Florida

geological history have only recently been developed. Geologic and

geophysical investigations for petroleum resources have provided the

primary subsurface information upon which these hypotheses are based. A

limited number of deep wells have extended into the Paleozoic basement

rocks.

Based on deep well data, Applin (1951) defined the contact between two

major rock types in the basement complex (Figure 2.5-3). The approximate

zone of contact forms a line which extends northeasterly across Florida

from about Tampa Bay on the west coast to about halfway between

Jacksonville and St. Augustine on the east coast. Northwest of this

contact zone are unmetamorphosed Paleozoic sediments; to the southeast are

Paleozoic granites and volcanics.

Aeromagnetic surveys conducted by Gough (1967) in the eastern Gulf of

Mexico show magnetic anomaly trends in Central Florida which tend to

substantiate the Applin Boundary. This contact zone between Paleozoic

sediments and Paleozoic igneous rocks may represent the southern limit of a

structural basin, or a fault boundary (Vernon, 1951; Applin, 1965;

2.5-2 Rodgers, 1970). There is no evidence of displacements of the overlying formations.

The surface configuration of the basement is undulatory, featuring relative highs (archs) and lows (basins). The depth to the basement complex varies

from about 2,600 feet in north central Florida to greater than 15,000 feet

in south Florida (Vernon, 1970). In the St. Lucie County area it is at a

depth of about 13,000 feet (Milton, 1972). The pre-Mesozoic structures

which produced this undulatory surface are discussed in Section

2.5.1.1.3.1.

Late Paleozoic and early Mesozoic strata have not been identified in Florida, due to non-deposition or erosion. Overlying the basement complex

are formations ranging in age from Jurassic through Recent. Applin and

Applin (Vernon, 1964) indicate that: "....submergence of the Floridan

Peninsula apparently began early in the Mesozoic Era...Jurassic(?) and

Cretaceous sediments were deposited in a transgressive sea which encroached

northward accompanied by progressive subsidence of the Coastal Plain

floor." This subsidence is generally responsible for both the dip and

general thickening of Mesozoic and Cenozoic formations away from the

relatively stable Peninsular Arch. The formations are primarily shallow-

water marine carbonate deposits, principally limestones and dolomites.

Unconformities exist between some of the formations due to intermittent sea

level fluctuations and resulting periods of erosion and non-deposition.

Within the Cenozoic Formations (primarily Eocene, Oligocene, and Miocene Formations) gentle warping (folding) has created localized areas of

thickening and thinning sequences of sediments. Examples of variations of

bedding thicknesses can be seen on the attached seismic reflection profiles (Figures 2.5-95 through 2.5-107). The time of warping varies from Eocene to

middle Miocene, as evidenced in the relationship of the bedding in the

Hawthorne Formation (Miocene) to the bedding in the underlying Suwannee and

Ocala limestones (Oligocene and Eocene).

During the interglacial stages of the Pleistocene Epoch, the sea level was higher for a part of the time than at present and portions of Florida were

submerged. Cooke (1945) postulated the existence of seven terraces which

correlate with different levels of the sea during the Pleistocene Epoch.

These can be more or less identified and correlated across north Florida, Georgia and the Carolinas.

During glacial stages of the Pleistocene Epoch, the sea level was lower than at present and some currently submerged portions of Florida stood

above sea level. Evidence for this exists in the presence of a buried

Pleistocene river channel discovered near the course of seismic reflection

profiling in the St. Lucie site area. The buried channel is an incised

river channel offshore of the town of Wabasso, just north of Vero Beach, Florida approximately 25 miles north of the site. The channels runs

generally southeast and has eroded through most of the Hawthorne Formation.

Offshore, the depth of this channel reaches approximately 500 feet beneath

present sea level. As the Pleistocene channel approaches the present

coastline, it becomes shallower. Reflection profiles crossing the channel

in the Indian River show it to be approximately 200 feet below present sea

level. This feature represents a major drainage feature during

Pleistocene time which has since become filled with late Pleistocene and 2.5-3 Amendment No. 25 (04/12)

Recent sediments. No trace of this river channel can be seen at the surface. 2.5.1.1.3 Regional Structural Setting of Peninsular Florida

Hypotheses with regard to the structure of the basement complex which underlies the Coastal Plain deposits in peninsular Florida are based on

limited data - samples obtained from petroleum exploration wells and

geophysical investigations; primarily aerial magnetic and gravimetric

surveys. These data indicate that the basement complex is an undulatory

surface, however data is insufficient to determine if elevation

differentials are the result of warping of the rocks which comprise the

basement or of differential erosion. Regardless of their mode of

formation, the Peninsular Arch and the basing around it appear to be

structures formed or initiated in pre-Mesozoic time. There is no

structural activity hypothesized to have occurred in peninsular Florida

during the Mesozoic and Cenozoic Eras. Mesozoic and Cenozoic sedimentary

strata were deposited in Florida, with their distribution and lithologies

controlled to a major degree by the pre-Mesozoic basement topography.

Vernon (1951) has hypothesized that the Ocala Uplift and associated local minor structures and faulting were produced in late Oligocene or early

Miocene times by compressive forces acting from the west against the

Peninsular Arch. These features are shown on Figure 2.5-3. Others have

hypothesized that these features, if they exist, are the result of

differential erosion, deposition and consolidation, principally during the

Eocene-Miocene period.

Vernon (1951) also recognized numerous joints or fractures in Florida.

Studies made by others, including Bermes (1958) have not shown a

relationship between Vernon's joint patterns and faulting.

The structural features discussed above are presented in more detail in the following sections. 2.5.1.1.3.1 Major Tectonic Structure

A tectonic structure is considered to be a deformation of major scale in the earth's crust produced by major force systems in the mantle. It is a

prominent regional feature. The major tectonic structures in peninsular

Florida are the Peninsular Arch, the Southeast Georgia Embayment, and the

South Florida Basin (Murray, 1961; Cohee, 1962). These features are shown

in Figure 2.5-3. These major tectonic structures are attributed to warping

or displacement of the basement surface prior to the Mesozoic Era.

The Peninsular Arch is a northwest - southeast trending subsurface high located in northern Florida and southern Georgia. It forms the axis of

peninsular Florida and has influenced the subsequent deposition of Mesozoic

and Cenozoic sediments.

The presence of the Arch was originally postulated by Applin and Applin (1944) based on a study of wells drilled through Coastal Plain deposits

into the basement complex. Most deep wells have terminated upon

encountering the basement complex, therefore, information regarding the

structure 2.5-4 of the Arch is limited. The core of the Arch is composed of early Paleozoic rocks. Applin (1951) describes the arch as an anticlinal fold.

Murray (1961) shows the Paleozoic rocks to be relatively undisturbed (flat

lying) and unmetamorphosed, inferring that the Arch is an erosional

feature.

Regardless of the origin of the Peninsular Arch, evidence indicates that:.

a) The Arch is a pre-Mesozoic structure comprised of Paleozoic rocks

b) Late Paleozoic and early Mesozoic rocks are not present due either to non-deposition or erosion

c) The Arch was a topographic high which was not covered over until the late Cretaceous, and

d) The presence of the Arch continued to influence deposition the Cenozoic period

The Southeast Georgia Embayment is located in coastal Florida, Georgia, and

South Carolina, northeast of the Peninsular Arch. The Embayment is

described as a shallow asymmetrical syncline, steeper on its southwestern

flank and opening to the east (Murray, 1961). Along the Georgia coastline, it contains about 6,000 feet of Mesozoic and Cenozoic deposits. The

thickness of the sedimentary sequence increases off-shore. Paleozoic, and

possibly some Precambrian, sediments and crystalline rocks underlie the

Coastal Plain deposits, forming the basement of the feature (Herrick and

Vorhis, 1963).

The South Florida Basin is located southwest of the Peninsular Arch.

Limited well data in south Florida indicates the axis of this structure is

located off-shore in the Gulf of Mexico with a northwest trend, plunging

northwest (Murray, 1961). The full sequence of Mesozoic and Cenozoic

deposits have not been fully penetrated by a maximum 15,000 foot deep in

south Florida. Paleozoic sediments and crystalline rocks apparently form

the basin basement.

2.5.1.1.3.2 Minor Structure

The minor structures in peninsular Florida are of Tertiary age. They are

the Ocala Uplift and associated Eocene-Miocene minor structures and local

faults as shown on Figure 2.5-3. The closure or displacement of these

structures is generally on the order of about 100 feet, with a maximum of

about 400 feet, versus several thousands of feet for the tectonic

structures previously discussed.

The Ocala Uplift is generally discussed as a northwest-southeast trending

gently folded structure centered in northeast peninsular Florida. The

highest point along its axis is located in Citrus and Levy Counties. In

north central Florida along its axis, the uplift is about 230 miles long

and 70 miles wide.

Two different hypotheses on the origin of late Tertiary structure in

Florida are postulated by Vernon (1951) and Brooks (1974). Vernon indi-

2.5-5 cafes that wells drilled through the central axis of the Ocala Uplift penetrate the flanks of the Peninsular Arch west of the axis of the arch.

He attributes the formation of the Ocala Uplift to late Tertiary forces, acting from the west against the Peninsular Arch. Brooks indicates that

the Ocala Uplift is an erosional feature.

Vernon (1951) also hypothesized the existence of minor structures and local normal faults of late Eocene-early Miocene age in Florida. He attributes

their development to the same forces which produced the Ocala Uplift. Most

of the faults are oriented parallel to the crest of the Ocala Uplift or are

postulated to form the boundaries of the other minor structures (Kissimmee

Faulted Flexure, Sanford High, and Osceola Low). Vertical offsets on these

faults are reported to range from less than 50 feet to up to about 400

feet, based on well cutting studies.

Other investigators have postulated the existence of additional late Eocene-early Miocene normal faults similar in origin to those described by

Vernon. Among these are four faults in Indian River and Martin counties.

Seismic reflection profiles made during the St. Lucie investigation

indicates structure consists of warping rather than faulting. This is

discussed in detail in Section 2.5.1.2.3.

Brooks (1974) has drawn contours on the top of the Ocala limestone which indicate that the Ocala Uplift is an erosional feature, probably produced

as a result of natural erosion processes acting on a topographic high

during Eocene-Miocene time. The minor structures and local faults

associated with this feature, if they exist, are probably the result of

differential consolidation created by gravitational forces acting on

varying thicknesses of the sedimentary sequence.

Of these two different hypotheses, that of Brooks appears the more reasonable. There is no direct evidence of the forces that would be

necessary, to produce the flexure postulated by Vernon. Further, because

of the Ocala - Miocene interval, non-conforming erosion would be logical. 2.5.1.1.3.3 Joints

Vernon (1951) recognized a system of fractures or joints which he indicates can be traced from county to county throughout the state. He uses the

terms "fractures" and "joints" interchangeably. The joints were mapped

based on physiographic features (predominantly lineaments of stream and

sink holes) recognized on aerial photographic mosaics. Most joints appear

aligned in a northwest-southeast or northeast-southwest pattern. A

secondary north-south, east-west pattern is also evident.

Vernon attributes the origin of the joints to structural movements during the late Tertiary. Specifically, he believes the joints are tension cracks

developed parallel and perpendicular to the compressive axis of the Ocala

Uplift. Straley (1968) has reported evidence for similar joints or

fractures in the southern portion of the state. He associates these joints

with the predominant structural trends in Florida.

2.5-6 An alternative interpretation for the development of the joint patterns is differential consolidation of relatively unconsolidated sediments as they

were deposited in the structural lows or draped over structural highs of

the stable basement complex.

Of the hundreds of "joints" mapped by Vernon in Florida, approximately 10

occur at locations coinciding with hypothesized faults. All of these occur

in central and northern Florida, over 100 miles from the site. A geologic

study was made by Bermes (1958) in Indian River County subsequent to

Vernon's 1951 report. Bermes study did not show any relationship between

Vernon's joint patterns and faulting.

2.5.1.1.4 Stratigraphic and Lithologic Setting of Peninsular Florida

Paleozoic igneous and sedimentary rocks form the basement complex in

Florida. The igneous rocks range from granites to basalt flows and

pyroclastics. The sediments are unmetamorphosed or weakly metamorphosed, relatively flat-lying, noncalcareous shales and sandstones (Applin and

Applin in Vernon, 1970). The sediments range in age from early Ordovician

to middle (?) Devonian.

The Paleozoic strata are overlain in peninsular Florida by from 2600 to

over 15,000 feet of Mesozoic and Cenozoic sediments. The thinnest sequence

of post-Paleozoic sediments is located in north central Florida, the

thickest in south Florida. These strata form a seaward thickening wedge of

southeastward gently dipping formations which extend off of the Piedmont

physiographic province to form the Atlantic and Gulf Coastal Plains.

The Mesozoic strata in northern peninsular Florida are a mixed facies of

clastics and silty and sandy carbonate rocks. To the south, they are

predominantly carbonates (limestone and dolomites) with some evaporites.

Cenozoic rocks in Florida up through the Oligocene period, are basically, shallow-water marine carbonates and occur throughout most of peninsular

Florida. Chen (1965) refers to this environment of deposition as the

Florida Platform. Paleocene and lower and middle Eocene sediments are

generally thick strata of dolomites with interbedded limestones. Late

Eocene and Oligocene formations are mostly fossiliferous limestones.

The overlying Miocene strata contains some basal limestones, however, in

most of Florida the Miocene is composed primarily of partially cemented and

indurated sands and clays. The cementing agent is predominantly calcite.

Pliocene deposits are generally partially cemented sands, shells

and limestones.

A large portion of peninsular Florida is covered by a series of Pleistocene

marine terraces comprised primarily of sands and shell fragments. The

terraces were most likely deposited during interglacial stages when the sea

level was higher than at present. Coastal currents transported the

basically quartzitic sands which comprise the terraces down into Florida

from the north along the Atlantic coast. The terraces are less distinct on

the west coast of Florida, possibly due to the lack of source sand

materials which were transported into Florida by Atlantic coastal currents.

In south Florida, sandy limestones and calcareous sands were deposited

2.5-7 rather than terraces, indicating a lower land surface (deeper water environment) in this area. This indicates a continued influence of

subsidence in south Florida through the Pleistocene epoch.

Figure 2.5-4 is a regional surface geology map showing outcroppings of Eocene to Recent deposits in Peninsular Florida. Figure 2.5-5 shows the

top of the Avon Park Formation of Eocene age. A geologic section (Figure

2.5-6) shows the relatively flat lying Coastal Plain formations between

Tampa and the St. Lucie site, and the influence of the basement Peninsular

Arch on these sediments. 2.5.1.2Subregional Geology A specific geologic investigation has been made of the vicinity of the

site, primarily Indian River, St. Lucie, and Martin Counties. This

investigation included air photo studies, evaluation of existing well data, and geologic boring and continuous seismic reflection profiles.

Particular emphasis was placed on defining stratigraphy in order to

evaluate a previously hypothesized structure in the area and the validity

of the extension of other structural features to points near the site.

2.5.1.2.1 Physiography An investigation of land forms (shown on Figure 2.5-1) has been made using satellite photography and U.S.G.S. Topographic Quadrangle Maps. Figure

2.5-7 is a print made from a satellite photograph made in 1973. An overlay

of Figure 2.5-7 (Figure 2.5-7A) shows land form subdivisions with

delineations based primarily on topographic and drainage characteristics.

The site is located within the coastal lowlands. The land forms along this part of the east coast of Florida generally parallel the northwesterly

orientation of the present coastline. This alignment of land forms

reflects depositional processes similar to those which exist at the

present. Recognizable land forms within the Coastal Lowlands are: 1) the

present coastline depositional environment, 2) the Atlantic Coastal Ridge, and 3) a broad very shallow valley or swale which contains the headwaters

of the St. John's River and the St. Lucie River. West of the Coastal

Lowlands is a higher terrace of the interior plain and valleys which

extends westward to the highland ridges.

The eastern-most land form within the coastal lowlands depositional environment includes the barrier islands and the lagoonal areas of the

Indian River. The barrier islands, including Hutchinson Island, were

probably formed as off-shore bars during a period of higher sea level.

These islands vary in width from a few hundred feet to about one mile.

Surface elevations on the barrier island are between sea level and about

+15 feet, with the higher elevations along the present coastline. The

western portions of the islands are primarily mangrove swamps.

The Indian River averages about two miles in width and is shallow (about 3

to 6 feet deep) except for a ten-foot deep dredged channel for the Inland

Waterway.

2.5-8 The western bank of the Indian River is the Atlantic Coastal Ridge with surface elevations ranging up to a maximum of about +40 feet. The ridge is

an almost continuous land form extending from the Sebastian River in

northern Indian River County south to the St. Lucie River at Stuart, in

northern Martin County. The average width of the ridge is about 1/4 to 1/2

mile. West of the Atlantic Coastal Ridge is a broad essentially flat valley, which is approximately 25 miles wide. Surface elevations within the valley

from southern Brevard County through Martin County generally range between

+20 and +30. Originally the valley was probably a lagoon, similar to the

environment of the present Indian River. It was later inundated by the sea

and presently reflects Pleistocene deposition of the Pamlico Terrace.

Studies of satellite photography and topographic maps indicate some

remnants, through subdued, of swale and swell topography. These features

suggest progradational beach ridges. However, relief is minor and most

topographic expression has been lost, possibly due to leaching over a long

period of time of the calcareous shell content of the sands (White, 1970).

There is a remnant of the Talbot Terrace in Indian River County, Figure

2.5-7A. Surface elevations on this terrace are about +40 feet. The

terrace shows several northerly trending beach ridges.

West of the Talbot Terrace and the valley within the Coastal Lowlands is a higher terrace (Penholoway) at a level with surface elevations of about

+60 to +70. The Penholoway terrace is the eastern portion of the major

land form classification, interior plains and valleys. It is approximately

40 miles wide and the topography tends to rise to the north. The western

boundary of this terrace is the Lake Wales Ridge, a north-south trending

ridge of the Central Highlands, located about 80 to 90 miles inland from

the present Atlantic coastline.

The eastern and southern portions of the Penholoway Terrace exhibit some

distinctive beach ridges and inter-ridge swales. Where these features are

present, they control drainage patterns. Inland of the beach ridges, the

terrace exhibits little topographic relief and drainage is more random, except for the northwest-southeast trending Kissimmee River. The series of

beach ridges along the eastern portions of the terrace probably mark an old

coastline which may have created a barrier separating an interior lagoon

from the ocean environment. White (1970), suggests that the Kissimmee

River may have developed as a drainage feature of this interior lagoon as

the sea level dropped. This mode of land form development is similar to

that suggested in the valley immediately inland of the Atlantic Coastal

Ridge. In essence, the geomorphic features characterizing the east coast of southcentral Florida are postulated to be a series of beach ridges and

inland lagoons, and environment of deposition not dissimilar to the present

relationship between the barrier islands, including Hutchinson Island, and

the Indian River.

2.5-9 2.5.1.2.2 Lithology and Stratigraphy

About 13,000 feet of Jurassic to Recent sedimentary formations overlie the

Paleozoic crystalline basement complex in the sub-region (Figure 2.5-1).

This thickness of sedimentary formations is based upon data from well

number W-4086. The location of this well is shown on Figure 2.5-8 along

with other well data points in the site area used for this report. The

upper 600 feet of sediments are soft rock formations consisting of

partially cemented and indurated sands and clays. Below 600 feet, these

materials are primarily moderately hard to hard limestones and dolomites

with some sandstones, shales, and anhydrates (Milton, 1972).

Lower Cretaceous and possible Jurassic formations extend upward from the

basement to an elevation of about -7,000 feet. The Jurassic and lower

Cretaceous formations pinch out against the flanks of the Peninsular Arch

and dip in a southeasterly direction. These formations consist of

limestones and anhydrites with some beds of clastic materials (Vernon, 1964). The clastic materials are most likely derived from erosion of the

rock composing the Peninsular Arch.

In St. Lucie County, upper Cretaceous materials extend from about -7000 to

-4200 feet (-4240 feet in well no. W-4086). The upper Cretaceous (Gulf

Series) in southern Florida is basically thick sequences of carbonates (often non-fossiliferous, chalky) with some sandstones and shales. Beds of

upper Cretaceous age also dip to the southeast in the site vicinity and

thicken away from the crest of the Peninsular Arch.

Cenozoic formations representing Paleocene through Recent age underlie the

site. The Paleocene and lower and middle Eocene formations are basically

hard, crystalline dolomites with interbedded fossiliferous limestones. In

well number W-4086 (located approximately 7 miles west of the St. Lucie

site) these formations were encountered at the following elevations: Cedar

Keys limestone (Paleocene), -3200 to -4200 feet; Oldsmar limestone (Paleocene), -2240 to -3200 feet; Lake City limestone (Eocene) -1340 to

2240 feet; and Avon Park limestone (Eocene) -800 to -1340 feet.

The upper Eocene in St. Lucie County is represented by the Ocala Group

limestone. This formation is basically soft to moderately hard

foraminiferal limestone. It is about 150 feet thick in well number W-4086 (elevations : -800 to -650 feet) and thickens to about 200 feet along the

coastline. The large foraminifera, Lepidocyclina, is a diagnostic indicator of this formation.

Overlying the Eocene sequence, is the Suwannee Limestone of Oligocene age.

The Suwannee varies from a hard, very fossiliferous, "coquinoid", limestone

to a soft, granular, dolomitic, essentially non-fossiliferous limestone

containing broken fragments of pectons, echinoids and barnacles. The lower

Suwannee sequence appears to grade into the Ocala sequence.

The Suwannee averages about 135 feet thick in borings drilled by Law

Engineering during its investigation of the site area. AG-104, AG-105, and

AG-106 are located in St. Lucie County immediately west of the site as

shown on Figure 2.5-9. Studies in Indian River and Martin Counties

2.5-10 indicate a thickening of the Suwannee to the east accompanied by an increase in dip. In northern and western Indian River County, the Suwannee

is absent. The Suwannee Formation is considered to be the uppermost

formation of the Floridan Aquifer in this region. The overlying Hawthorne

Formation is a major confining unit (aquiclude) for the Floridan Aquifer.

The Hawthorne Formation of Miocene age unconformably overlies the Suwannee Limestone. This formation averages about 430 feet thick in borings AG-104, AG-105, and AG-106 located on the mainland west of the site. The lithology

of the Hawthorne is basically green, clayey, very fine sands and very fine

sandy clays. Phosphatic sand and clays and thin layers of sandy, phosphatic

limestone are characteristic of the lower part of the formation. The

formation is generally dense and indurated.

The contact between the Hawthorne Formation and the overlying Anastasia Formation is unconformable. Prior to deposition of the Anastasia Formation

the Hawthorne was subjected to a period of erosion which left its surface

essentially flat in the vicinity of the site, as shown in the geologic

sections on Figures 2.5-9, 2.5-10, and 2.5-11.

The Anastasia Formation (contemporaneous) in the site area refers to all the material which overlies the Hawthorne Formation except for Pleistocene

Terraces. The formation is of Pleistocene age. It is about 140' thick at

the coast and thins to the northwest against older more steeply dippings

formations. To the southwest it either pinches out or merges with the Ft.

Thompson Formation. The Anastasia differs in composition from place to

place, ranging from practically pure quartz to almost pure quartz sand.

2.5-11 There are generally thin discontinuous pockets and seams of sandy limestones or sandstones within the sequence, An example of these scattered

discontinuous layers is exposed along the coast at St. Lucie Inlet 15 miles

south of the site. The Anastasia is the main source of water supply for

several towns and cities in this coastal area, including Fort Pierce and

Stuart.

The Anastasia Formation is overlain in some areas by Pleistocene marine

terraces. These terraces are composed of quartzitic sands and shell

fragments, The Atlantic Coastal Ridge, which forms the western bank of the

Indian River is a remnant of such a terrace.

2.5.1.2.3 Structure

No specific reports on the geology and groundwater resources of St. Lucie

County have been published; however, geologic studies have been reported

for Martin County, to the south, and Indian River and Brevard Counties

to the north, Figure 2.5-12 is a map showing all hypothesized structures

in the sub-region.

Bermes (1958) utilized data from wells drilled into Eocene strata to de-

velop east-west and north-south geologic sections in Indian River County.

These sections indicated the Eocene and Miocene strata sloped gently to the

southeast in most of the county. Bermes reported an apparent change in dip

from less than 5 feet per mile to greater than 70 feet per mile and the

occurrence of Oligocene age beds near the eastern margin of the county. He

postulates a somewhat complex system of three high-angle, normal faults

essentially parallel to the coastline to explain the steepening dip and the

occurrence and apparent thickening in Oligocene strata to the east. Strata

on the east side of the faults were projected to be downthrown. The faults

were inferred from a difference in elevation of about 225 feet of tho top

of the Ocala Group for a horizontal difference in elevation of about 2.5

miles between the control points. The age of the faulting was not discussed

by Bermes, but the fault traces shown in his geologic sections were

terminated at the bale of Miocene strata, indicating an age greater than 20

million years.

Lichtler (1960) utilized electric logs and some cuttings from deep wells in

a study of Martin County. Based on this study, Lichtler postulated a high

angle northwesterly trending normal fault which is parallel to and about to

5 miles inland of the present coastline. The strata on the east side of the

fault were projected as being downthrown by as much as 300 to 400 feet (Lichtler, 1960). The maximum displacement between control points (an

electric log of well 443 and a study of cuttings from well 841) is a

displacement of 270 feet for a horizontal distance of about 2.5 miles (Lichtler, 1960). Lichtler dated the faulting as offsetting strata at least

as young as the Oligocene with some possible additional slippage during the

early Miocene, He attributed the taulting to the forces which formed the

Ocala Uplift, thus establishing an age of over 20 million years.

Vernon (1970) used hydrologic data to extend the Lichtler fault in Martin

County through St. Lucie County (partly along the St. Lucie River) to a

point at about the southern boundary of Indian River County. The shortest

distance from this hypothesized fault to the site is 5 miles. His control

2.5-12 was the elevation of the top of the artesian aquifer (Floridan Aquifer) and not any particular geologic unit.

Our study of the regional geologic structure, particularly in-the vicinity

of the St. Lucie site, included verification of existing well data

previously used by Bermes and Lichtler in their hypotheses of area

structure drilling deep geologic borings, the performance of approximately

50 miles of marine continuous seismic reflection survey, and meetings with

Bermes, Lichtler, and other geologists familiar with the St. Lucie area.

Data from a number of previously drilled water wells to the south, west, and north of the site (including key data points used by Lichtler and

Bermes in their reports on Martin and Indian River counties) were

incorporated into the study. Prior to utilization of this data, samples

from these wells on file at the Florida Geological Survey were studied.

Nine geologic borings were drilled outside of the immediate plant area on

Hutchinson Island. The borings were drilled north, south, and west of the

plant site at distances of from two to six miles. The locations of

previously existing borings and additional geologic borings drilled by Law

Engineering are shown on Figure 2.5-8.

The primary purpose of six of these borings was to verify the elevation of

the top of the Hawthorne Formation at depths consistent with those found at

and adjacent to the plant site. Figures 2.5-9, 2.5-10, and 2.5-11 show the

top of the Hawthorne to be relatively flat surface, sloping slightly down

to the east at about 3 feet per mile, as would be expected from the

regional structure.

Three deep geologic borings (700 feet +) were drilled along a line west of

the site to obtain data on both sides of the Vernon "fault" and on both

sides of the hypothetical straight line extension of the bermes "fault" which would pass closest to the site (about 3 miles west of the site).

The location, spacing, and depth of the three deep geologic borings were

dictated by the nature of the structure postulated by Vernon, and a

straight line extension of the structure hypothesized by Bermes. Since

these hypothesized structures are based on scattered well data several

miles apart, the published locations of the structures are accurate only to

within the limits of the distances between borings upon which they are

based. Thus, the three borings drilled to investigate the structures were

placed on two mile centers perpendicular to the hypothesized structure to

insure that the structure was adequately bracketed. Since the postulated

structures were based primarily on differences to top of limestone (top of

Suwannee Formation) in scattered wells, the geologic borings were extended

far enough into limestone such that specific geologic formations could be

identified and correlated.

Within the depths drilled, three formational contacts were identified by

lithology, electric and gamma-ray logs, and fossils. These contacts are

the top of the Hawthorne Formation, top of the Suwannee Limestone, and the

top of the Ocala Group Limestone. These are formations which have been

hypothesized in the published literature to have been offset by faulting.

2.5-13 Figure 2.5-9 is a geologic section drawn through the borings where they cross the hypothesized Vernon fault and the closest extension of any of the

Bermes faults to the plant site. This section shows flat lying strata with

insignificant elevation variations across proposed or hypothetical faults.

It is concluded that no faulting exists at these locations. Our

interpretation of 2 marker horizons (top of Ocala Group and Avon Park

Limestone) are shown on Figures 2.5-13 and 2.5-14.

In order to more closely define structure in the vicinity of the St. Lucie

site, and to further evaluate the existence of faults hypothesized in the

literature, approximately 50 miles of continuous marine seismic reflection

profiles were made. The survey was accomplished using a maximum 3000 joule

sparker energy source firing at 1/1 second intervals and continuously

monitored and recorded in strip chart format. This reflection surveying

system was capable of detecting marker horizons within the upper 1000 feet

of materials. The key marker horizons were correlated with deep boring data

described.

As a part of the survey, the fault system hypothesized by Bermes in Indian

River County was crossed at two locations where the structure trends across

the Indian River into the Atlantic. The fault hypothesis by Lichtler was

crossed at a location on the north fork of the St. Lucie river just north

of the Martin County - St. Lucie County Line. Additionally, reflection

profiling was accomplished immediately north, south, east, and west of the

site, as well as adjacent to the Unit 1 reactor.

Within the depths penetrated, several key marker beds were identified. The

most significant and the strongest reflecting horizon represents top of

limestone (top of Suwannee Formation). Since differences in elevation of

top of limestone as determined from scattered well data are the primary

basis for postulated faults, the characteristics of this key horizon were

studied in detail along the profile line locations. Generalized profile

line locations are shown on Figure 2.5-92. Detailed line locations, showing

navigational control points, are shown on Figures 2.5-93, 2.5-94, and 2.5-

95. Sections of individual profiles are shown on Figures 2.5-96 through

2.5-110.

Three prominent seismic reflectors are present throughout the survey area.

These have been correlated to drill hole AG-104, and represent the top of

the Suwannee limestone (top of limestone in the site vicinity), and two

distinct layers in the Hawthorne Formation. In addition, the top of the

Hawthorne Formation was detected in a number of areas covered by the

survey, as was the top of the Ocala limestone. All layers throughout the

survey area are predominantly flat lying, with regional dip to the

southeast being evidenced best by the top of the Suwannee reflector, which

rises from approximately 750 feet near St. Lucie Inlet to about 300 feet

Sebastian Inlet in Indian River County to the north.

Superimposed on the regional dip of about 10 feet per mile, are localized

areas of warping (folding) in which dips on the order of 150 feet per mile

or about 3 degrees exist. This type of warping was found to exist in Ft.

Pierce Inlet, St. Lucie Inlet, in Indian River county near Sebastian

2.5-14 Amendment No. 25 (04/12)

Inlet, and underlying the St. Lucie site.

These four areas illustrate the effects of localized warping on depth to top of limestone over relatively short horizontal distances. Under the Fort

Pierce Inlet, a downwarp occurs as shown on Figure 2.5-95. Folding of the

sediments has created a syncline with a closure of 150 feet in the limestone over a horizontal distance of approximately one mile (1.7dip).The profile of St. Lucie Inlet (Figure 2.5-101) shows a similar synclinal downwarp with closure on the order of 200 feet over a horizontal distance of approximately two miles (1.4dip). In the area just south of Sebastian Inlet, a monoclinal downwarp occurs as shown on Figures 2.5-104, 2.5-105, and 2.5-106. In this area, the top of limestone slopes downward from

elevation -200 to nearly elevation -700 feet at a point two miles to the east (2.9dip). Figure 2.5-98 shows a monoclinal downwarp under big Mud Creek adjacent to the site with closure on the order of 150 feet over a

horizontal distance of one mile (1.7dip).Local variations in depth to the top of the limestone (Suwannee reflector) evidently was the basis for the hypothesized faults in Martin County and

Indian River County. In Indian River County, one of the faults hypothesized

by Bermes, and which showed apparent offsets in limestone on the order of

225 feet, was investigated by reflection profile lines crossing between

three key boring locations. This area is described above as an area of

warping near Sebastian Inlet. The reflection profile shows locally strong

dips of up to 250 feet per mile in top of limestone. This variation in top

of limestone is in agreement with logs of the three borings used by Bermes.

However, the reflection profiles show continuous beds with no evidence of

faulting.

The reflection profile made across Vernon's extention of Lichtler's fault in Martin County provides only sketchy data of subsurface structure in this

area. However, the data which was obtained, as shown on Figure 2.5-109, has

the same pattern of reflectors as that found throughout the seismic

reflection survey. The decrease in elevation in the top of limestone to the

southeast is typical of the regional dip, and shows an elevation decrease

in top of limestone of about 100 feet eastward over a horizontal distance

of approximately three miles. This is consistent with variations in

elevation of the top of limestone found in the adjacent St. Lucie Inlet. It

is concluded that faulting does not exist at this location.

Following collection of the data described above, meetings were held with Bermes and Lichtler to discuss the correlation and significance of this

additional data relative to previously published structure. Both Bermes and

Lichtler agree that no significant offsets exist anywhere along the

approximately 50 miles of reflection profiles.

In summary, a series of roughly parallel layers in a fairly constant pattern with respect to depth has been traced from St. Lucie Inlet south of

the St. Lucie plant, past and adjacent to the plant site, and north to

Sebastian Inlet, approximately 30 miles north of the plant site. No faults

of any kind were found in the sediment sequence. Several areas of localized

and possibly connected warping were found, with most of these located under

or near the present barrier island. Sediments under the Indian River (and

at least 6 miles inland from the immediate site vicinity as determined 2.5-15 Amendment No. 25 (04/12) by the deep geologic borings) are nearly flat, as are those of the offshore area.

Time of warping apparently took place from the Eocene to middle Miocene

Epochs, as evidenced in the various relationships of the bedding in the

Miocene Hawthorne formation to the limestone surface below. There does not

appear to be any evidence for warping at the present time, as the seismic

reflector identified as top of the Miocene (Hawthorne) appeared to be

nearly horizontal over the entire survey area.

Utilizing all of the data obtained, it is concluded that faulting does not

exist within the upper formations in the vicinity of the St. Lucie plant.

2.5.1.2.4 Potential for Surface Subsidence

2.5.1.2.4.1 Solution Activity in Carbonate Terraines

Solutioning of carbonate rocks and resulting karst topography is well

developed in some portions of Florida. However, a study of satellite

photography shows no evidence of advanced solutioning of carbonate bedrock

formations and resulting lake development within 50 miles of the site.

White (1958, 1970) has discussed solutioning in Florida and its mechanisms

of development in detail. The site vicinity is not recognized as a high

potential area for surface subsidence due to solutioning. Three factors

which appear to be limiting influences on solutioning in this area are 1)

the young age of the coastal land forms relative to the geologic time

required to produce mature karst, 2) the depth to carbonate bedrock and 3)

the thickness and clayey nature of the Miocene sediments which overlie the

limestones.

White (1970) indicates that a principal mechanism for solution development

is initiated in beach-ridge covered terraces where water is concentrated in

the inter-ridge swales. These land form features are evident in terraces 20

miles west of the site where beach ridges, though subdued, are abundant.

2.5.1.2.4.2 Mineral Extraction

Several local communities obtain their water supplies from the Anastasia

Formation. In order to assess the potential effect of water extraction on

surface subsidence, officials of the two largest municipal users nearest

the site, Fort Pierce and Stuart, were contacted.

According to Mr. C. W. Temby, Director of the Fort Pierce Utility

Authority, Fort Pierce pumps a maximum of 10 MGD from about 30 wells

installed to a depth of 120 to 130 feet. They project a usage of 24 MGD by

1980. Additional wells will be installed to the same depth as those in

current use to provide this additional capacity.

In Stuart, Mr. D. A. Gale, Water and Sewer Superintendent, stated that

Stuart pumps an average of about 2.5 MGD from 22 wells, 120 feet deep.

2.5-16 Additional wells will be installed to increase capacity to 6 MGD.

Monitoring stations in the Stuart well field record minimal drawdown with

maximum pumpage,

Both Messrs Temby and Gale state that there is no record of subsidence

related to well field pumping in either Stuart or Fort Pierce. Since there

is no subsidence associated with these two well fields, and since the St.

Lucie site is separated from communities utilizing water from the Anastasia

Formation by the Indian River (an infinite source of water), there is no

anticipation of subsidence at the site related to ground water withdrawal.

Other than ground water, there is no mineral potential known in the site

area. Several oil test wells have been drilled in St. Lucie, Martin, and

surrounding counties. Most of these wells were terminated in the basement

rock complex. No oil or gas "shows" were recorded in any of these wells, and all have been capped and abandoned.

2.5.1.3 Site Geology

2.5.1.3.1 Geologic History

The site is located on the East Coast of Florida on an offshore bar named

Hutchinson Island, as shown on Figure 2.5-1. The East Coast of Florida is

the emergent coast of Florida. Hutchinson Island has been developed and has

remained exposed above MSL since some time during the Pleistocene age, approximately one million years ago. The Pleistocene sediments are slightly

thicker at the site than at the mainland due to the method by which the

island was developed.

2.5.1.3.2 Physiography

The topography of Hutchinson Island is a bar and inland swale type. There

is a low bar near elevation +14 on the ocean side of the island. The

surface of the island then slopes downward toward the Indian River to about

elevation +4, generally forming a swale. To the north and west of the site, both Big Mud Creek and the Indian River are continuations of the swale and

are very shallow, 5 to 10 feet deep. There is a dredged channel in the

Indian River for the inland waterway. To the east, the Atlantic Ocean

bottom slopes very slightly to the east to a depth of about 120 feet at a

distance of about 15 to 20 miles from Hutchinson Island.

To the west, at the mainland, there is another bar (Atlantic Coastal Ridge)

with a maximum elevation of about +40 parallel to the coast.

The site was originally covered by a mangrove swamp. A small road dike, for

mosquito control, was constructed around the portion of the site that is

adjacent to big Mud Creek and the Indian River. Within the site area at the

time of initial field work, there was about 1 foot of standing saline

water.

2.5-17 2.5.1.3.3 Stratigraphy and Lithology

There was 4 to 6 feet of peat and roots beneath the generally water covered

surface. This material is a dark brown or black residuum produced by the

partial decomposition and disintegration of trees, mangrove roots, and

other vegetation. This peat was probably formed during the past several

thousand years.

The Anastasia Formation of Pleistocene age underlies the peat. No

differentiation has been made between the various Pleistocene deposits

which extend to about elevation -135 to -155 feet below sea level. This

material has been termed "Anastasia Contemporaneous". There is a suggestion

from geologic and engineering evidence that the discontinuous cemented

pockets are the erosional remnants of the Fort Thompson Formation. If this

is true, then generally below elevation -35 to -60 are the Anastasia and

Caloosahatchee Formations. The Anastasia Formation is quite possibly

composed of marine, brackish and fresh water deposits. Petrographic studies

have shown that at some places there is well rounded and frosted sand

indicating a beach or marine depositional environment. At other places

within the formation there is a blackening of shells which may indicate a

reducing, or brackish depositional environment. Also identified was a fresh

water gastropod from one of the partially cemented, discontinuous pockets,

The Anastasia Formation consists of gray, slightly clayey and silty, fine

to medium sand with fragmented shells and, in places, fragmented shell beds

with slightly clayey and silty, fine sands. There are also discontinuous

pockets of cemented sand with shells and sandy limestone. These

discontinuous cemented pockets are generally found from about elevation -35

to -60. Also occasional discontinuous plastic clay lenses were found in the

upper part of the formation.

The Hawthorne Formation of Miocene age unconformably underlies the

Anastasia Formation. There is evidence of this unconformity or erosional

surface by possible "case hardening" of the contact of the Hawthorne and

Anastasia. This occurs between elevations -140 and -157.

The upper 100 to 150 feet of the Hawthorne Formation consists of green, slightly clayey and silty, very fine sand. The lower part becomes generally

more clayey. Geologic information obtained by seismic reflection pro files

in Big Mud Creek show the Hawthorne formation as extending downward to

about elevation -600 to -700 in this area.

This lithology of the Hawthorne Formation changes slightly to a gray white, phosphatic, sandy clay at this site below possibly elevation -450. The

clays of the Hawthorne are unique in their content of the mineral

Attapulgite. The x-ray diffraction analyses also indicate a high carbonate

fraction (calcite and dolomite) in the Hawthorne Formation.

2.5.1.3.4 Structure

The structure of the geologic formations at this site is simple, as

determined by borings (Figure 2.5-15) and reflection profiles (Figure 2.5-

93). The younger, or Pleistocene and Miocene formations are nearly flat, dipping very slightly to the southeast at about 5 to 10 feet per mile.

There is an

2.5-18 erosional surface or unconformity between the "Anastasia Contemporaneous" and the underlying Hawthorne formation. This has resulted in a slightly

undulatory contact having minor irregularities (Figures 2.5-16 and 2.5-17).

The older or deeper formations (limestone formations) dip at somewhat

greater angles. Seismic reflection profiles made adjacent to the site show

dips increasing with depth, At a depth of approximately 800 feet, the

limestone formations are dipping 150 feet per mile toward the southeast.

This is consistant with regional and area structure,

2.5-19

Borings 1 through 20 as shown on Figure 2.5-23 were drilled in 1968 prior to site preparation. Cross sections through these borings are shown on

Figures 2.5-24, 25, 26 and 27. Seven of these borings were performed in

the plant area proper and the remaining thirteen borings were performed in

the general site area. These borings were drilled to evaluate the general

site geologic conditions and to provide preliminary subsurface information

on static and dynamic soils conditions. Geophysical surveys were performed

at that time.

In early 1969, Borings 101 through 146 were drilled. Twenty-seven of these

additional borings were located in the immediate plant area to supplement

the data obtained from the initial site boring program. The remainder were

drilled at locations of site related facilities. The purposes of the

supplementary foundation investigations were as follows:

a) To verify in further detail the subsurface conditions at the location of the structures

b) To establish the liquefaction resistance of the soils in greater detail c) To develop additional foundation data to verify bearing capacity and estimated settlements

d) To determine the physical properties of the compacted soils for use in the dynamic evaluation of the soil foundation structural system

e) To locate suitable fill material

Additionally, at this time auger borings A-1 through A-4 were performed to

obtain samples for laboratory investigation of the proposed fill.

Borings 147 through 173 and SB-1 through SB-5 were performed as continuing

studies of site related facilities, such as ocean intake and discharge

pipelines, bridge pier foundations (Highway A-1-A) and switchyard.

Additional geophysical surveys were performed in early 1973 to obtain shear

wave velocities of materials within the construction area.

Graphical descriptions of the soils encountered in the main borings are

shown on the logs of borings presented in Appendix 2A. Geophysical survey

test results are presented in Section 2.5.4.4. The field and laboratory

testing procedures used in this investigation are discussed in Sections

2.5.4.3 and 2.5.4.4. Laboratory test data is presented in Appendices A and

C for the St. Lucie Unit 1, PSAR.

2.5-31 2.5.4.3 Properties of Underlying Material Samples for laboratory testing were obtained from the test borings in order

that the proper ties of the underlying materials could be determined. The

borings were advanced by a rotary drilling process which utilizes a viscous

bentonite drilling fluid to flush the cuttings and stabilize the hole. At

regular intervals, the drilling tools were withdrawn and soil samples

obtained with a standard 1.4" I.D., 2.0" O.D., split-tube sampler. The

sampler was initially seated 6" to penetrate any loose cuttings then driven

an additional foot with blows of a 140 lb hammer falling 30". The number

of hammer blows required to drive the sampler the final foot was recorded

and is designated the "Standard Penetration Resistance".

The samples, as they were obtained, were classified in the field by an

engineering geologist. Portions of each soil sample were sealed in glass

jars and transported to our laboratory where they were examined by soils

engineers and geologists.

Core drilling was performed generally in accordance with specification ASTM

D 2113-70. Prior to initiating coring operations, 4-inch I.D. casing was

installed to a sufficient depth to prevent caving of overburden soils into

the hole. Boring was performed using an NX or HQ, double-tubed, swivel

type core barrel. Upon completion of each run, the core was removed from

the barrel and logged by a geologist. All core was carefully placed in

wooden boxes and wrapped in plastic to prevent excessive moisture loss.

Undisturbed samples were taken of selected strata for laboratory testing.

The samples were obtained by forcing 30" long sections of 3" O.D.,

stainless steel tubing into the soil. The sampling procedures were

conducted in accordance with ASTM Procedure D 1587. The samples thus

secured, were sealed with paraffin to prevent moisture loss and transported

to the laboratory.

Bore hole logging of borings AG 104, AG 105, and AG 106 was performed upon

completion of drilling. Logging was performed by personnel of the Division

of Water Resources, State of Florida. A Gearhart-Owens Model 3200 portable

logging truck mounted system was used to perform the work. Data was

recorded on a Gearhart-Owens MRP 501 x-y recorder.

A single point electric probe was used to simultaneously record self

potential and electrical resistivity. A gamma-ray record was also obtained

for each boring.

The properties of underlying materials are identified with the major

geologic formations encountered. The two major formations encountered and

their general engineering characteristics are discussed below.

The Anastasia formation occurs at the surface of the plant site and varies

in thickness from 135 to 155 feet. The Anastasia formation consists of

grey slightly clayey and silty fine to medium sand with fragmented shells

and, in places, fragmented shell beds with slightly clayey and silty fine

sands. There are also discontinuous pockets of cemented sand with shells

and sandy limestone. These discontinuous cemented pockets are generally

2.5-32 found from about elevation (-) 35 to (-) 60, except at Boring B-1 (Boring logs are in Appendix 2A) where they extend to about elevation (-) 90. Also

occasional discontinuous thin plastic clay lenses were found in the upper

part of the formation.

The Hawthorne formation underlies the Anastasia formation. The upper 100

to 150 feet of the Hawthorne formation consists of a green slightly clayey

and silty very fine sand. The lower part becomes generally more clayey.

The published geologic information describes the Hawthorne formation as

extending downward to about elevation (-) 600 in this area. This was

partially substantiated by our deep geophysical exploration at the site.

The Hawthorne formation was found to extend to a depth of at least 600

feet; no harder layers were encountered to this depth. The Hawthorne

formation changes slightly to a grey white sandy clay at this site below

possible elevation (-)450.

This formation is generally dense and indicated by both desiccation and

cementation.

The borings indicated that the soil could be separated into three zones

depending upon compactness. The first or Upper Zone 50 to 60 feet was a

loose sand with small amounts of silt and clay, containing isolated pockets

of shell fragments and limestone nodules.

As described in Section 2.5.4.8 the studies have shown that the Upper Zone

was potentially subject to liquefaction. Therefore, this zone required

remedial treatment beneath all critical structures as described in Section

2.5.4.5.

Beneath non-critical structures, the upper sands provide support for light

structures. They have moderate strength and are relatively incompressible.

Heavier structures could cause consolidation of the clayey lenses, and

analysis is described in Sections 2.5.4.10. The recent organic or peat mat

underlying the Upper Zone was unsuitable for foundation support, and

therefore was removed.

An Intermediate Zone extends from about 60 feet to 150 feet in depth. The

soil of the Intermediate Zone differs from the shallower soil in that it is

denser, contains a greater percentage of fines (material finer than the

number 200 sieve), and has very few pockets of limestone nodules and shell

fragments. This zone is differentiated from the Upper Zone on the basis of

consistency and grain size characteristics.

In this zone beneath the plant there do not appear to be any isolated

pockets of limestone nodules and shell fragments. This finding coupled

with the increased amount of fines in the zone from 60 feet to about 150

feet, separates it from the Upper Zone. The material in the Intermediate

Zone has properties which will not adversely affect the foundation. It is

strong enough to support the loads produced by surcharging and by the

imposed weight of the structures.

2.5-33 The Deep Zone extends from 150 feet in depth to at least 400 feet. This material is considerably more clayey than the material above, does not

contain pockets of shells and limestone, and is dense. The soils are

normally consolidated under the existing overburden load.

Table 2.5-2 shows penetration resistance and percent fines of certain

borings.

At depths greater than 150 feet the standard penetration test does not give

reliable values. The consistency of this material was determined from

triaxial test results, from torvane shear values, and from an approximate

correlation of shear strength with the force required to push a standard

split-spoon sampler into cohesive soils. The results of these tests are

given in Table 2.5-4. This latter method was utilized since the drill rod

and hammer weight was greater than 1000 pounds and significantly influenced

the results of the standard penetration test. The approximate method of

determining the shear strength of cohesive soil using a split-spoon sampler

is shown in Figure 2.5-28. The soil consistencies were determined from the

following table:

CONSISTENCY OF CLAY IN TERMS OF UNCONFINED COMPRESSION STRENGTH (Reference 30)

Unconfined Compressive Strength Consistency (kg/cm

2)

Very soft Less than 0.25 Soft 0.25 - 0.5 Medium (firm) 0.5 - 1.0 Stiff 1.0 - 2.0 Very Stiff 2.0 - 4.0 Hard Over 4.0

In all cases, the consistencies were verified by visual examination. From

analysis of laboratory test data it is concluded that this deep material

does not present any foundation stability problem. It is anticipated that

the minor amount of consolidation settlement will be such that the

consolidation phase of the settlement is expected to be complete during the

construction operation, or shortly thereafter.

Samples extracted from the borings were subjected to laboratory tests to

determine the physical properties of the soils. Summaries of the

laboratory data are given in Appendix 2A. Laboratory test data is

presented in Appendices A and C for the St. Lucie Unit 1 PSAR. The

complete laboratory program included the following tests:

a) Grain size analysis b) Specific gravity c) Moisture content and density d) Maximum-minimum density e) Proctor compaction test

f) Consolidation.

2.5-34 Additional laboratory tests were performed to determine the static and dynamic shear properties of the soils. These tests include the following: a)Direct shear b)Triaxial compression c)Cyclic triaxial shear d)Compression wave velocity.

Direct Shear and Triaxial Compression tests were performed on

representative samples of the compacted material that was used for the

plant backfill. These tests further established the soil parameters, such

as the angle of internal friction and the effect of pore water pressure, which are utilized in bearing capacity evaluations.

2.5.4.4 Soil and Rock Characteristics In order to determine the seismic effects on the plant structural foundations, the geophysical properties of the site were examined. The

scope and mode of this examination is presented below: a)Determination of Geophysical Parameters for Seismic Analysis of Structures A refraction seismic exploration was undertaken at Hutchinson Island to

determine the compression wave velocities under dynamic conditions and also

to determine if there were any significant changes in the rigidity of the

formations (which could be detected by seismic velocity) below the deepest

borings made at the site. The strains that accompany this type of field

test are very small, i.e., microstrains. The corresponding compression

wave velocity represents an upper limit. The wave velocity for larger

strains, such as those induced by strong motion earthquake, would be

smaller. An additional use of the data from the geophysical exploration

was to determine if there were any significant differences in elevation

between the lithologic boundaries that might suggest structural

displacements, in a wider area than that covered by the onsite borings.

Twenty-eight refraction lines of varing lengths were made in the site vicinity. The lines were spaced around the perimeter of the site, along

State Road A1A to the east, and along the mosquito control dike south and

west of the site.

In general, the seismic refraction profiles indicated there were no significant variations in the elevations of the boundaries between the

strata at the St. Lucie site. A few of the seismic lines indicated

materials with velocities of approximately 6660 feet per second exist

between the surface and 150 feet. These higher velocity zones correlate

with local concentrations of limestone nodules and cemented sand shell

lenses found in the borings.

The instrument used was a Dresser Model RS-4, 12 channel, recording seismograph with surface geophones capable of recording compression waves.

The charges of explosives detonated ranged from 2 to 15 pounds of gelatin

dynamite placed both at the ground surface and in bore holes as deep as 70

feet. 2.5-35 In several instances, it was possible from the seismic refraction work to calculate velocities in the two principal formations underlying the site;

the upper sands of the Anastasia Formation (to a depth of about 150 feet),

and the deeper clayey Hawthorne Formation. In traverses where compression

wave velocity could be differentiated between these two formations, it was

found to be 5700 feet per second between 50 feet and 150 feet, and 6800

feet per second below a depth of approximately 150 feet. These two

velocities are very similar, making a delineation of the Anastasia-

Hawthorne contact difficult on the basis of the seismic refraction work

alone. However, in those instances where velocity differences were

obtained, the Anastasia-Hawthorne interface depth was found to be in

agreement with borings drilled in the vicinity.

If it is assumed that each material is a homogeneous elastic mass, it is

possible to calculate the elastic modulus from the density of this material

and the compression wave velocity. Using density data available from the

borings at the site, typical total densities were used in calculating the

approximate dynamic modulus of elasticity. The upper limit elastic modulus

for the sands of the Anastasia Formation was calculated to be approximately

210,000 psi, and for the deeper Hawthorne Formation, approximately 285,000

psi.

Due to the varying densities it was not possible to determine with

reliability the compression wave velocity in the uppermost loose material

above a depth of 50 feet. Further the water table was approximately at the

ground surface. In saturated materials whose wave velocity in the dry

state is less than the velocity of water, the indicated wave velocity will

be that of the water in the voids rather than of the formation itself.

Thus, based on rigid control conditions for the backfilling operation (assuming uniform density) the elastic properties for the compacted fill

foundation were calculated as follows:

Four cyclic triaxial compression tests were conducted to establish the

compressional wave velocity of the compacted backfill. The tests were

conducted on samples compacted to 85 percent relative density and at

residual confining pressures developed during compaction in order to

simulate the ground surface conditions. Crystal transducers were used for

sending and receiving the ultrasonic waves, thus the induced strains were

in the microstrain region as is the case with the field seismic work. Nine

tests were made varying moisture content and percent fines and all tests

confirmed a compressional wave velocity of 700 fps with results not varying

more than 5 percent. Other investigators verify these results (9,33).

Assuming that the materials are a homogeneous elastic mass and utilizing

the measured compressional wave velocity and a Poisson's ratio of 0.25 as

determined from the above references, the near ground surface elastic

modulus was calculated to be 12,000 psi. It has been shown by other

investigators, including the ones referred to above, that the elastic

modulus increases with depth and is approximately proportional to the

square root of the depth. Thus at the level of the reported field test the

calculated upper bound elastic modulus from laboratory data will be

approximately twelve times the surface value or about 150,000 psi.

2.5-36 The comparison of the elastic modulus as measured in situ and as determined in the laboratory is reasonable. If consideration is given to the in-place

soil formation and structure as compared to a remolded structure of a

compacted soil, the comparison is quite favorable. This phenomenon of

original soil structure or in-situ structure has long been realized to have

an important effect on the characteristic of soils, in particular the

elastic properties.

Since the strain conditions developed during earthquake shocks are much

greater than the very small strains induced to determine the elastic

properties as given above, consideration must be given to selecting

parameters to be used in the elastic and shear moduli and Poisson's ratio.

Evaluation of the elastic moduli was carried out in two steps. First the

modulus applicable to very small strains (microstrains) was determined.

This is usually done by utilizing field or laboratory seismic tests as

described above. Then reduction factors were used to adjust the modulus

for the expected level of strain.

Once strains exceed the 10

-5 to 10-4 level and non-linear effects become noticeable, there is no single universally accepted definition of velocity

or modulus. As the peak strain involved in a repeated loading application

increases, the modulus and velocity decrease. Some investigators have

expressed this decrease by means of a reduction factor, which relates the

ratio of the velocity for a level of strain to the velocity for very small

strains (10). Another investigator, H.B. Seed (11) states that "...a change in strains from the magnitude associated with microtremors to the magnitudes

associated with major earthquake may cause a 2 to 10 fold decrease in

modulus... while microtremor effects can serve an extremely useful purpose

in establishing one bound on the range of possible behavior patterns and in

checking the applicability of a proposed analytical procedure, they do not

appear to provide, in themselves, a full predictive capability for

engineering purposes."

In order to evaluate the effects of the magnitude of the elastic constants

on the dynamic analysis of the Class I structures, both the numerically

larger modulus as determined from the field seismic work, and the smaller

modulus as determined from the laboratory studies, reduced as described

above, were considered. This approach served to establish an upper bound

modulus, determined from microstrains, and a lower bound modulus, determined by reduction factors described above.

It was determined that the numerically large elastic constants would yield

a condition in the range of the peak of the response spectra. Therefore, any other values of the elastic properties could only yield numerically

smaller response characteristics. Field measurements indicated that the

elastic modulus was 210,000 psi and laboratory measurements of the

compacted fill indicated a ground surface elastic modulus of 12,000 psi.

Preliminary, separate dynamic analyses of the Reactor Building structure

were run using elastic moduli ranging from 10,000 psi to 250,000 psi and it

was found that the maximum response acceleration occurred at an elastic

modulus of approximately 150,000 psi. These preliminary maximum

acceleration values were used for the design of the structures. When the

design

2.5-37 was completed the structures were again analyzed using the elastic moduli which resulted in the preliminary maximum acceleration. If a significant

difference in acceleration resulted from the re-analysis, a range of

elastic moduli were run again to determine the most conservative values.

For the sake of conservatism, the numerically larger elastic moduli were

utilized to determine the spring constants in the structural dynamic

analysis. This approach of using the modulus determined from microstrain

seismic methods is conservative. The method of seismic analysis of Class I

structures is discussed fully in Section 3.7.

b) Determination of Geophysical Parameters for Seismic Analysis of Equipment

When the parametric values which were used for seismic Class I structure

analysis were used for determining the floor response spectra envelopes for

seismic Class I equipment analysis, the design of piping systems became

unrealistic because of the wide range of periods for which peak resonant

accelerations were indicated by the spectral curves. Therefore a further

soils testing program was initiated to determine a more realistic value of

soils modulus by determining the relationship with respect to strain rather

than utilizing the parametric range for the building analysis.

The following laboratory program was set up to determine the applicable

soils modulus.

1) Existing cyclic triaxial tests were studied to establish the moduli and strain relationships. Originally the data was not reduced to

determine any stress-strain functions.

2) Existing computer studies of soil response had developed moduli and corresponding strains. These were reviewed in order to determine the

compatability of the strains to the strains developed as a result of

these additional investigations.

3) Additional cyclic tests were performed at stress and strain levels of interest using soil samples obtained from the field under proper

relative densities and confining pressures.

4) The developed relationship of soils modulus versus strain was plotted and analyzed and is presented as Figure 2.5-32.

Two hundred pounds of sand were obtained from the site. Initially, a grain

size test was performed on a representative part of the sample. This test

was performed in accordance with ASTM standards.

Maximum and minimum density tests were also performed on a representative

portion of the sample. The maximum density was determined utilizing a wet

sample on the vibrating table. The minimum density was determined by

pouring dry sand through a funnel into a known volume. A specific gravity

test was also performed on a portion of the sand sample in accordance with

ASTM D 854. The results of the specific gravity and density determinations

were utilized for maximum and minimum void ratio calculations, which are:

2.5-38 Void Ratio, maximum - 0.987 minimum - 0.609 Dry Unit Weight, maximum - 104.0 pcf minimum - 84.2 pcf Specific Gravity - 2.68 Cyclic Triaxial Shear Tests were run on samples 2.8 inches in diameter by

5.6 inches

long that were compacted to 85 percent relative density. The

samples were then placed in a cyclic triaxial shear chamber and subjected

to a confining pressure of 3,500 psf. Water was allowed access to the

sample and backpressure of approximately 50 psi was applied to insure

saturation. During the backpressure saturation, the effective confining

pressure was maintained at 3,500 psf. Cyclic deviator stresses varying

from 180 psf to 1,700 psf were applied with a bellowfram loading system at

a frequency of 2 cycles per second. During cyclic loading, continuous

record of load and deformation were monitored by calibrated strain gauge

devices and recorded. Both load and deformation monitoring devices were

mounted within the triaxial chamber.

The dynamic modulus of elasticity, "E", was calculated from the peaks of the load and deformation outputs. In effect, the calculated moduli are the

secant values from peak to peak of the hysteresis loop. The dynamic shear

modulus, "G", was calculated using the relationship: ) + 2(1 E =G µPoisson's ratio, "" was assumed to be 0.45 for saturated sand.

For low ranges of strains for which the concepts from the theory of

elasticity are of use, Poisson's ratio () varies with strain. The value has been shown by other investigators to be near 0.5 for very low strain

values, on the order of 10

-6 and to decrease to near 0.2 for strain levels near 10-3 and then increase to values above 0.5 at very high strains for dense sands. It is difficult to make an exact evaluation of the value of

"" for use in solution to real problems; but for dense soils even the range from 0.5 to 0.2 has a relatively small effect upon engineering

calculations and evaluations. The 0.45 value was utilized in this section

to evaluate the value of the dynamic shear modulus of elasticity at strains

of 10-5 in/in. This assumed value of "" has since been confirmed for this strain level using the values of the field compressional and shear wave

velocities obtained from the cross hole technique discussed in Section

2.5.1.4(c), using the following relationship: 1-)V/V (1-)V/V 1/2( = 2sp 2spµwhere: V P = Compressional wave velocity 2.5-39 V S = Shear wave velocity

Table 2.5-3 presents a tabulation of deviator stress, strain and moduli.

Figure 2.5-11 is a graph of shear modulus versus shear strain.

The shear modulus of sand is dependent upon several variables. Among the

most important variables, which are not functions of the soil properties, are the effective confining pressure and the range of strain to which the

sand is subjected. Seed and Idress (11) have published a tentative relationship for a medium sand (D r 75 percent) which accounts for both confining pressure and strain. When this data is extended to an effective

confining pressure of 3,500 psf, close agreement occurs with the data

developed by this study. At strains above 3 x 10

-4 inches per inch the Hutchinson Island data is higher than the published data. This is

partially explained in a somewhat stiffer material. Slight fluctuations in

pore water pressure may also have influenced the Hutchinson Island tests

but were not accounted for. At any rate, the data for Hutchinson Island

agrees very well with the data presented by Seed and Idriss.

The tests for Hutchinson Island represent a single level of confining

stress. The confining pressure used was selected as the average effective

overburden pressure at the level of the Reactor Building foundation. In

order to extend this data to other levels, the variation in confining

pressure should be accounted for. The modulus of elasticity was found to

vary exponentially with the confining pressure. Sowers and Sowers (12) state that:

E = C3 n where "n" varies from .35 to .82 depending on the principal stress ratio, Seed and Idriss state:

G = k 33 n where n = .33

As previously mentioned, the test results for Hutchinson Island sand agree

well with the data presented by Seed and Idriss, particularly in the strain

range of 1 to 2 x 10

-4 inches/inch. Therefore, it is concluded that the relationship developed by Seed and Idriss can be utilized to extend the

Hutchinson Island data to confining stresses other than 3.5 ksf (24.3 psi).

That is:

24.3 G - G 3x 3.5 x

Where:

G x = Shear modulus desired at confining pressure

G 3.5 = Shear modulus from Figure 2.5-11 3x = Confining pressure at level G is desired, (psi)

2.5-40

)-()-( = Dd

The dynamic test equipment and the procedures used were similar to those described by Seed and Lee (31).

In the older version of the cyclic triaxial machine the pneumatic device

applies an approximate square shaped wave excitation to the test sample.

Because of this, the peak load is not instantaneous but applied for a

finite duration of time. The transition to the peak load is not smooth; in

addition to this, the inertia of the complete system becomes significant

and the load is applied in a non-uniform way. Since this sharp build-up and

holding of load existed the best loading conditions were not met, and

sample failure was premature when compared to a smoothly applied sinusoidal

excitation.

The newer versions of the cyclic triaxial loading mechanism eliminates most

of these limitations. The hydraulic system programmed applies a smoothly

varying sinusoidal excitation where the peak load is gradually developed, held as an instantaneous peak and not applied for a finite time causing

premature failure. In addition to the above conditions, the load pattern

would be erratic and might begin at the desired deviator stress but, has in

the past, due to equipment friction and inertia, been seen to vary up and

even down during continued cycling. The load dropping off during continued

cycling could lead to unconservative results at a high number of cycles of

higher strain levels.

Data from Ebasco's experience with soils and recent cyclic triaxial tests

were assembled and reviewed. Figure 81 shows the effect of equipment and

procedures on cycles to produce initial liquefaction (i.e., Pore Pressure =

Confining Pressure) and clearly shows the conservatism of older equipment

and procedures. It is concluded that in spite of many variables coming into

play in a liquefaction analysis, data conclusively prove the conservatism (lower stress ratio dp/2a causing liquefaction of the older equipment and testing procedures. In fact, the increase in stress ratio at 10 cycles

using today's techniques would be equal to 1.78 the ratio of

0.32 (obtained from Figure No. 81)).

0.18

The soil sample was subjected to a uniform confining stress approximately

equivalent to the overburden stress at the depth of concern. The sample was

then subjected to cycles of axial loading. Tests were run at axial load

magnitudes both larger and smaller than those that might be induced by

earthquakes at the site. The cyclical or pulsating axial loads were

maintained as constant as possible with a frequency of approximately one

cycle per second (corresponding to the longer shear pulses of a strong

motion earthquake) until the sample failed.

Failure generally occurred in two stages. The first indication of impending

failure was an increase of the pore water pressure which momentarily

equalled the confining stress on the soil. However, with the reversal of

load the pore water pressure changed and the soil exhibited little change

in resistance through the continuing load cycle. This "momentary

liquefaction" is defined by the number of load cycles required to produce a

momentary pore water pressure equal to the confining stress. This does not, however, mean that substantial permanent strains would occur in the soil.

2.5-51 In these tests, momentary liquefaction generally occurred at double amplitude strain levels between 0.5 percent and 3 percent, with one

exception at 5 percent. These strain levels varied depending upon the

relative density confining pressure and peak pulsating deviation stress.

With additional cycles of load, the duration of the high pore water

pressure became progressively longer and the soil had little or no strength

through most of the load cycle. The strain increased rapidly, until it

became impossible to impose the full cyclic load on the sample.

At this point, the soil had lost its strength and liquefied. This point is

defined as "liquefaction" or by some, "complete liquefaction". Generally, in these tests, this corresponded to strains of about 15 percent. This

compares very well with the definition of "complete liquefaction" of 15

percent strain adopted by Seed and his co-workers (28) . The above definition of "momentary liquefaction" is more severe than Seed's definition of "initial liquefaction" at 5 percent strain.

From the test results, graphs were plotted depicting the shear stress at

the time of both momentary and complete liquefaction as a function of the

number of cycles of load required to produce liquefaction. Figure 2.5-44

shows the number of cycles required for momentary liquefaction; Figure 2.5-

45 shows liquefaction or complete liquefaction. The test data, including

the back pressure valves and "B" coefficient, is presented on Table 2.5-6.

The relative density was determined as follows.

A single maximum-minimum density determination was made on a composite

sample. The composite sample was obtained by combining shelby tube samples

obtained from boring B-106, 87 ft to 89 ft, and boring B-113, 87 ft to 89

ft and 97 ft to 99 ft. The maximum density of the sample was determined by

the modified proctor test (ASTM D1557 - Method C), and the minimum density

was determined by the dry funnel method such as described by ASTM 2049-64T (ASTM 2049-69). The results of the test are:

Maximum density 116.5 PCF Minimum void ratio 0.500 Minimum density 88.9 PCF Maximum void ratio 0.966

The weight-volume relationships were measured for each sample subjected to

cyclic triaxial shear tests. The relative densities, therefore, are the

relationship between maximum, minimum, and sample void ratio.

The scatter in the data is the result of using samples at different

relative densities. Straight lines are shown on Figures 2.5-44 and 2.5-45

as conservative lower bounds of the data, with respect to a minimum

relative density of 55 percent as discussed above. The data on page A-13 of

the Unit No. 1 PSAR was originally replaced with Figures 2.5-44 and 2.5-45

after a review of the original data indicated errors in the testing and in

the data on page A-13.

2.5-52

= The results of the tests are expressed in terms of the cyclic deviator stress divided by the effective confining stress '. This is based on the findings of Seed et al (31) that the cyclic shear stress at failure is proportional to the effective confining stress. Therefore, tests run at one

confining pressure can be used to analyze failures at other confining

stresses. Seed has expressed the shear stress at failure, induced by 10 cycles of load by the expression: = 'R d/200 In this expression:

= Shear stress for liquefaction in 10 cycles

' = Effective confining stress

R d = Relative density

For cycles different than 10 the value of 200 would be replaced by a con-

stant.

The shear stress is half the deviator stress so the expression could also

be rewritten as follows for any given number of cycles:

or

This expression also includes the relative density. For the intermediate

ranges of relative density of more than about 30 percent and less than

about 70 percent, the shear resistance increases in proportion to the

relative density (31). Therefore, tests run at one relative density can be converted to other relative densities by this procedure.

The above relationship using a relative density of 85 percent was used to

develop the curve for the compacted fill in the liquefaction analysis shown

on Figure 2.5-46. The results of actual laboratory cyclic triaxial tests

used in the analysis are presented in Figure 2.5-55 and on Table 2.5-6.

The results of triaxial dynamic tests are somewhat overly optimistic in

assessing the dynamic shear strength of soils (32) . Based on the results of simple shear tests and correlation with observed liquefaction of soils, Seed has suggested that the realistic dynamic shear strength of soils is 55

percent of the strength by triaxial testing.

Figure 2.5-47 shows the effective shearing resistance of the undisturbed

soils with respect to liquefaction between El -60 and El -150 feet at the

site.

The shear stress imposed by the design basis earthquake at any level can be

found by the surface acceleration and the profile of acceleration versus

2.5-53 depth. This profile of the acceleration versus depth was developed during a comparative study of acceleration at Hutchinson Island and Green Cove

Springs. In general, the profiles at Hutchinson Island show a decrease of

acceleration with increasing depth. At a depth of 60 feet the acceleration

is 0.86 times the surface acceleration. At a depth of 100 feet it is 0.79

times the surface accelerations. At a depth of 150 feet it is 0.78 times

the surface acceleration.

The total overburden stress at any level is found from the weights of the soil in the overall area. These soil weights together with acceleration are

utilized to calculate the dynamic shear stress at the appropriate level.

These dynamic shear stresses are similarly plotted on Figure 2.5-47.

The margin of safety against liquefaction can be expressed by the ratio of dynamic shear resistance to the dynamic shear stress. The plots show that

for complete liquefaction the ratio ranges from 2.4 at a depth of 60 feet

to 1.7 at a depth of 150 feet. Such margins are considered more than

adequate against liquefaction. It should be emphasized that these margins

are based on test data on samples with 55 percent relative density, which

corresponds to the lowest 12 percent of the samples. Moreover, an

inspection of the boring records indicates that the possibility of

liquefaction in the deeper portions of the zone between -60 and -150 is

academic because of the scarcity of the looser zones within that range.

Finally, experience of liquefaction in actual earthquakes suggests that

liquefaction is confined to the upper soil layers and has never been

observed deeper than 100 feet.

Figures 2.5-46 and 2.5-47 are not comparable as presented since the shear stress induced in the soil by the DBE for Figure 2.5-46 was calculated using the relationship = h a max and uses a datum of elevation +18.0 g while Figure 2.5-47 uses a relationship of avg = (.65)h a max (Rd) g and a datum of elevation +0.0. Each was prepared at different time frames in accordance with the state of the art and are not meant to be compared.

Section 2.5.4.8.1 combines this data with depth using the equation avg = (.65)h a max (Rd) g In the above expressions: =shear stress induced by earthquake =the unit weight of the soil h=depth a max =maximum ground surface acceleration rd =depth stress reduction coefficient Rd =depth stress reduction coefficient determined during comparative study 2.5-54 g = 32.2 ft/sec 2 0.65 = converts to avg

Summary of Liquefaction Potential

a) Upper Zone

All of the data demonstrated that a part of the Upper Zone was potentially subject to liquefaction during earthquakes. The remedial

treatment is described in Section 2.5.4.5.

b) Intermediate Zone

Within the Intermediate Zone there are a few isolated points at which the penetration resistances are less than 15 blows per foot. In each

of these cases the percentage of fines is much greater than the

critical conditions based on the Niigata criteria. A more detailed

investigation was directed toward developing information to evaluate

the occurrence, character and significance of this area. Based on

this investigation and on the Niigata criteria, it was concluded that

the soils in the Intermediate Zone between about 55 and 150 ft are

not susceptible to liquefaction.

c) Deep Zone

The Deep Zone is clayey. Such materials are not susceptible to liquefaction.

2.5.4.8.1 Simplified Procedure for Evaluating Soil Liquefaction Potential - Applied to the St. Lucie Plant Site

To supplement the evaluation already presented and to be able to compare

results from El +18 ft to -150 ft, the liquefaction potential was also

analyzed on the basis of Seed and Idress' simplified method.

(40)

The position of the water level is important in the analysis of

liquefaction potential. Based on the information in Section 2.4, a water

level of +2.0 has been used in the calculation presented. This is the mean

high water level and is, therefore, a conservative water level to use in

the liquefaction analysis. This is the highest water level expected at the

site without assuming simultaneous improbable hydrologic conditions (e.g.,

due to hurricane or flood) with the DBE. However, any additional rise in

water level due to such events would be a surface phenomenon, since the

permeability of the soil would not noticeably raise the water level in the

soil beneath the plant.

For this analysis, the stress ratio causing liquefaction was determined

from cyclic loaded, dynamic triaxial tests conducted on representative

samples of granular soils. Figures 6 and 7 from Reference 40 are reproduced

and presented as Figure 2.5-48. Seed actually used data from the St. Lucie

site. These points are indicated on the respective curves. In

2.5-55 Figure 2.5-48 the test results are presented in terms of 50 percent relative densities. The relative densities applicable to the St. Lucie site

are tabulated below:

St. Lucie Soils Loosest Relative Average Relative Density Density Backfill 85 97 Natural Soils El. -60 to -100 55 85(+) Natural Soils El. -100 to -150 55 68(*) (+) Based on average standard penetration resistance, Figure 2.5-51.

(*) Based on average standard penetration resistance, Figure 2.5-52.

The data from Figure 2.5-48 must be adjusted with respect to relative density. The shear stress required to cause liquefaction is proportional to

the relative density. This relationship is:

()()x=where: ()= Stress ratio for sample at x relative density

()D dp a= Stress ratio for sample at 50 percent relative density dp = Cyclic deviator stress a = Triaxial confining pressure Using the lowest value from each graph on Figure 2.5-48, very conservative stress ratios were selected. As the number of significant stress cycles

resulting from the DBE (a max = 0.1g) would be less than 10, ten cycles of strong motion were used as well as 30 cycles of strong motion, assumed to

occur from a distant earthquake with a max = 0.05g. The stress ratio at

various depths causing liquefaction for each case was computed from the

following relationship:

()=where: = In situ effective shear stress causing liquefaction = Stress ratio from test program causing liquefaction 2.5-56 dp = Cyclic deviator stress a = Triaxial deviator stress 'o= In situ effective overburden stress Cr = A correction factor applied to laboratory triaxial test data to obtain stress conditions causing liquefaction in the field. The "Cr" correction accurately corrects for

the effects from laboratory to field conditions by

reducing the effective in situ shear strength. The

dynamic stress distribution within the soil mass beneath

and adjacent to the foundation structures is not

different than the stress conditions under which triaxial

testing or direct shear testing is usually carried out

since the porous stones at the top and bottom of the

laboratory samples simulate the foundation mat, and the

sides of the direct shear apparatus or the confining

membrane simulate rigid conditions similar to those

adjacent to the foundation structures.

To assess the liquefaction potential of the site soils, the shear stresses causing liquefaction were compared to the average shear stresses induced by

the DBE or the shear stresses induced by a distant earthquake causing a

maximum acceleration at the site of 0.05g and 30 cycles of strong motion.

Assuming that the site soil is a deformable body rather than a rigid body, the average shear stress induced in the soil was calculated by the

following relationship: [][]=where: avg= average shear stress =total unit weight of the soil (pcf) h =depth (ft) a max =maximum ground acceleration g =gravitational acceleration constant (32.2 ft/sec

2) rd =stress reduction factor accounting for soil acting as a deformable body0.65 = factor converting maximum shear stress to average shear stress The average uniform shear stress, avg , and the shear stress causing liquefaction, 1 , shown on Figures 2.5-53 and 2.5-54 were compared and the safety factor against liquefaction was calculated by dividing the shear 2.5-57 stresses required to cause liquefaction by the shear stresses developed during the postulated seismic event. The minimum safety factor against

liquefaction thus calculated assuming loose soil conditions as discussed in

Section 2.5.4.8 was 2.16 for a max = 0.1g and 10 cycles of strong motion.

Also assuming loose soil conditions, a max = 0.05g and 30 cycles of strong motion, the minimum safety factor against liquefaction was calculated to be

3.72. Also shown on Figures 2.5-53 and 2.5-54 are graphs comparing the data

for average soil conditions as discussed above. This is a more realistic

condition. The minimum safety factors from this data are 3.09 and 5.31.

These high safety factors again demonstrate that liquefaction would not

occur at the site under the postulated conditions.

It appears that the lowest safety factors occur at El -150 ft and that the

safety factor might actually be lower than this if calculations were

performed for deeper soil strata. This occurred because one of the

conservative assumptions used in the calculations assumed a constant

reduction factor, r d , below a depth of 100 ft while actually it still decreases at a reduces rate. However, including this further reduction

would increase the safety factors below 100 ft. In addition, no

liquefaction has ever been observed at these depths and is precluded by the

high confining pressures caused by the overburden soils.

2.5-58 2.5.4.9 Earthquake Design Basis The selection of the earthquake design basis is presented in Section

2.5.2.10.

With respect to a paper regarding seismic criteria for nuclear power plants (Coulter, Waldron, Devine, "Seismic and Geologic Siting Considerations for Nuclear Facilities," 5th World Conference on Earthquake Engineering, Rome, 1973), these authors point out the subjective nature of intensity-

acceleration relationships and provide one of the over 40 available

relations. In order to apply the relationship between intensity and

acceleration presented on Figure 1 in the Coulter, Waldron, Devine paper, the foundation conditions must be evaluated in accordance with general soil

mechanics principles as well as dynamic or wave propagation considerations, that is, a conventional evaluation, in soils and foundation terms, must be

made in order to place the site in one of the three categories, 1) below

average, 2) average or 3) above average. Once the general category is

determined it is a simple matter to obtain a ground acceleration.

Figure 2.5-33 presents the commonly accepted soil mechanic values for sand

materials. The soil values presented in this Figure were compiled using

data from the references listed. Relative to Figure 2.5-33 and using the

values listed below:

Relative Density 85% above EL-60 lowest 20 percentile 62% below EL-60 and conservative as discussed in sections 2.5.4.8 and 2.5.4.10

Angle of Internal Friction 40 above EL-60 conservative based on 34 below EL-60 test values presented in Appendix A&W Vol 4 for the St. Lucie Unit 1 PSAR (entitled Hutch-inson Island Plant Unit 1 PSAR)

Standard Penetration 23.2 above EL-50 conservative as shown (50 + in plant area) in Figs. 2.5-49, 50 45.5 above EL-100 and 51 43.3 above EL-150

Shear Modulus 3x10 6 psf conservative as shown on Fig 2.5-32

It is obvious that the St. Lucie site fits into the middle range of the

average zone. For an intensity VI MM, at the St. Lucie site, using the

Coulter Waldron and Devine intensity - acceleration relationship, a value

of 0.07 is obtained for the ground acceleration which is again below the

minimum requirements which governed the selection of 0.1g. We consider the

0.1g maximum surface acceleration to be a very conservative figure, when

compared with actual and seismic conditions in peninsular Florida.

2.5-59 Amendment No. 25 (04/12) 2.5.4.10 Static Analyses 2.5.4.10.1 Bearing Capacity

The ultimate bearing capacities of the compacted backfill foundation as

determined for the Class I structures by utilizing the Terzaghi and Peck

Bearing Capacity Equations (30) . The soil parameters used were those determined by laboratory tests; an angle of internal friction of 40 with the compacted backfill at 85 percent relative density.

The ultimate bearing capacity for the containment structure is that of a

circular foundation in a cohesionless soil and is calculated as follows (30): q ult. = 1 D f N q + 0.6 2 RN Where: 1 = The effective unit weight of overburden soils (lbs/cu ft)

2 = the effective unit weight of bearing soils (lbs/cu ft)

D f = The depth of the foundation base below grade (feet)

N q N = bearing capacity factors dependent upon the angle of internal friction

R = the radius of the foundation (feet)

q ult = (80.5 lb/ft 3 ) (43 ft) (60)+0.6 (60 lb/ft

3) (80.0ft) (90)

q ult = 207,700 + 259,200 = 466,900 lb/ft 2

This bearing value is extremely large and is that soil pressure necessary

to cause a shear failure of this large diameter foundation resting on and

embedded in, 43 feet of dense compacted soil. The utilization of soil

pressure of even one-tenth of the above calculated value would yield

intolerable settlements; however, the concept is presented to illustrate

the fact that the utilization of allowable bearing values in the range of

10,000 to 12,000 lb/ft 2 are extremely conservative with respect to foundation bearing failure or instability. The criteria which govern the

allowable bearing capacities were total and relative settlements among the

various component plant structures.

The dynamic bearing capacity of the foundation materials was analyzed by

subjected Class I structures to oscillations due to hurricane buffeting.

This analysis considered the following: (1) the design hurricane winds of

130 mph sustained with gusting to 195 mph and reducing to 65 mph. (2) the

frequency of one cycle every one to three minutes and a duration of 30

hours.

On the basis of the above criteria the imposed foundation level loadings

2.5-60 Amendment No. 25 (04/12)

+/-

I Mc A P+=

°

==

i i W i W W W S F t n ST===

==°°==i S F ST

()()()()i cos 0.1 i sin tan i sin 0.1 i cos i Cos W 0.1 i sin W tan i sin W 0.1 i cos W+=////

°)

TABLE 2.5-1 HAWTHORNE CLAY X-RAY DIFFRACTION ANALYSES Total Sample (Per Cent)

B-1 B-1 B-1 B-1 Constituents 271 Feet 285 Feet 310 Feet 365 Feet Quartz 65 30 45 20 Feldspar 10 15 8 7 Calcite-15 24-Dolomite 20 20 23 55 Opal - Phosphorite- - -5 Clay 5 15 Trace 13 Clay Fraction (Per Cent)

B-1 B-1 B-1 Constituents 271 Feet 285 Feet 365 Feet Illite 30 30 25 Montmorillonite 30 35 30 Kaolinite 25 30 25 Chlorite 15 5-Attapulgite

-- 20 2.5-79 TABLE 2.5-1A ST. LUCIE UNIT NO. 1 SETTLEMENTS

STRUCTURE CALCULATED ACTUAL

Reactor Building 0.9 in. 1.1 in.

Reactor Auxiliary Building 0.5 in. 0.5 in.

Turbine Pedestal Mat 1.0 in. 1.32 in.

Intake Structure 0.0 in. 0.0 in.

2.5-80 TABLE 2.5-1B

SUMMARY

OF DYNAMIC SETTLEMENT TEST RESULTS 1.Sample was combined and remolded from undisturbed samples obtained from the Intermediate Zone2.Soil Description a) gray silty, fine, gravelly coarse to fine sand (fine gravel and sand sizes are shell fragments and sand) b)D 50 = 0.11 mmc)% Passing #200 Sieve = 17 3.Dry Density = 98 lb/ft 34.Moisture Content = 26%5.Relative Density = 55%

6.Confining Pressure = 6760 psf 7.Cyclic Stress Ratio = 0.181 8.Final Axial Strain = 0.1%

2.5-81 TABLE 2.5-2 PENETRATION RESISTANCE AND PERCENT FINES FOR BORINGS B-4,5,6,15,19,20 Standard Depth Penetration Boring No. (Feet) Resistance % < #200 Sieve B-4 60 34 13 65 81 12 70 42 13 75 30 8 80 100+ 10 85 41 7 90 57 12 95 95 30 100 97 25 B-5 5 23 6 7 22 4 10 21 1 15 35 1 20 19 2 23 18 2 30 14 29 36 11 63 47 10 3 70 - 10 125 - 23 128 - 10 130-6 135 50/3" 37 140 40 33 145 - 11 B-6 30 4 76 50 65 27 55 57 8 60 37 10 65 20 16 70 54 23 75 59 15 80-7 90 61 7 95 105 25 100 108 23 B-15 35 20 0 50 64 2 55 41 4 60 41 13 65 60 17 2.5-82

TABLE 2.5-2 (Cont'd)

Standard Depth Penetration Boring No. (Feet) Resistance % < #200 Sieve 70 64 12 76 45 18 80 86 13 85 100 13 90 83 23 95 75 43 100 81 45

B-19 70 40 17 75 12 22 80 19 28

B-20 70 89 18 75 16 17 80 62 26

2.5-83 TABLE 2.5-3 CYCLIC SHEAR TEST DATA

x 10-5 psi IN/IN d+/- +/- E psi SERIES I 0.60 1.32 45,000 1.72 4.06 42,300 2.76 6.70 41,400 4.41 11.45 38,400 5.70 17.40 32,700 6.30 23.20 27,300

SERIES II 0.38 (180) 1.26 2.66 47,400 3.32 7.10 46,800 4.41 11.55 38,100 7.71 26.65 28,800

SERIES III 6.79 18.6 36,600 9.15 32.3 28,500 (1,700) 11.98 40.4 29,700

2.5-84 TABLE 2.5-4 SHEAR STRENGTH

SUMMARY

BORING ELEVATION BY TORVANESHEAR STRENGTH (PSF)BY WEIGHT OFMATERIAL NUMBER (FEET) (WITHOUTBY TRIAXIALHAMMER AND RODDESCRIPTION CONFINEMENT)SHEAR TEST B-1 141 2800 Clayey Silty Fine Sand 230 6940 Silty Clayey Very Fine Sand 245 7310 Silty Clayey Very Fine Sand 250 700 3090 Clayey Very Fine Sandy Silt 255 1800 3140 Clayey Very Fine Sandy Silt 260 800 3200 Clayey Very Fine Sandy Silt 265 1400 3250 Clayey Very Fine Sandy Silt 270 1900 3300 Clayey Very Fine Sandy Silt 275 1600 3350 Clayey Very Fine Sandy Silt 280 1900+ Clayey Very Fine Sandy Silt 285 1800 Very Fine Sandy Clayey Silt 290 1000 Very Fine Sandy Clayey Silt 295 1900+ Very Fine Sandy Clayey Silt 300 1800 5100 Clayey Silty Very Fine Sand 305 5180 Clayey Silty Very Fine Sand 325 9310 Fine Micaceous Sandy Clayey Silt 340 5690 Very Fine Sandy Clayey Silt 345 9810 Very Fine Sandy Clayey Silt 350 9930 Very Fine Sandy Clayey Silt 355 10050 Very Fine Sandy Clayey Silt 360 5980 Clayey Very Fine Sandy Silt 365 6060 Clayey Very Fine Sandy Silt 380 10680 Very Fine Sandy Clayey Silt 385 10800 Very Fine Sandy Clayey Silt B-3 158 6500 Slightly Silty Fine Sand B-5 25 Fine Sand With Shells 37.5 1300 Slightly Fine Sandy Silty Clay 218 6560 Slightly Fine Sandy Silty Clay 223 6680 Slightly Fine Sandy Silty Clay 228 2830 Slightly Fine Sandy Silty Clay 233 2890 Slightly Fine Sandy Silty Clay B-6 30 1400 Slightly Fine Sandy Silt B-8 175 6000 Slightly Fine Sand 2.5-85 TABLE 2.5-5 REGIONAL EARTHQUAKE

SUMMARY

(HISTORIC EARTHQUAKES FELT IN PENINSULAR FLORIDA)

October 29, 1927: Several secondary sources report a severe earthquake in St.

Augustine on this date but the original record has not yet been located. New England suffered a severe shock about 10:40 A.M. on this date, and an earthquake was reported from Martinique on the same day (l).Distance of epicenter from site: 190 miles Epicentral Intensity: VI MM estimated February 8, 1843: Great earthquake centering at Guadeloupe, West Indies, N.

Latitude 16, W. Longitude 62, which did considerable damage and which appears to have been felt in the eastern part of the United States especially at

Washington, D.C.

(2). Distance of epicenter from site: 1500 miles Epicentral Intensity: Unknown

January 12, 1879: A general earthquake felt through north and central Florida from a line drawn from Punta Rassa to Daytona on the south, to one drawn from Tallahassee to Savannah on the north, an area of about 25,000 square miles (l).Epicentral location at N. Latitude 29.5, 14. Longitude 82.0 (2).Distance of epicenter from site: 200 miles

Epicentral Intensity: VI MM (2) January 22, 1880: Severe shocks (intensity VIII MM) were felt in Key West reflecting a disastrous earthquake in Vuelto Abajo, west of Havanna. This quake was felt and rumblings heard in Western Cuba and on the Isle of Pines.

It covered about 65,000 square miles (l). Epicentral location at N. Latitude 22.8, W. Longitude 80.8 (2). Distance of epicenter from site: 300 miles Epicentral Intensity: VIII MM (2) August 31, 1886: This is the date of the great earthquake at Charleston, S.

C., N. Latitude 32.9, W. Longitude 80.0. It was felt all over northern Florida, church bells rang in St. Augustine, and severe shocks were felt along

that part of the east coast. Apparently this quake had an intensity of V MM in

Florida (l). Distance of epicenter from site: 380 miles Epicentral Intensity: IX MM (3) September 1, 3, 5, 8, and 9, 1886: Jacksonville experienced more tremors of about intensity IV MM from the Charleston earthquake (l). Distance of epicenter from site: 380 miles 2.5-86 TABLE 2.5-5 (Cont'd)

October 22, 1886: Strong shocks at Charleston; the first in morning felt with intensity VI MM in Charleston, Atlanta, Augusta, and elsewhere, and second

with intensity VII MM at Sumerville, S. C. Felt at Washington Richmond, Louisville, Fayetteville, Jacksonville, and elsewhere (2). Felt at intensity V MM in Jacksonville (l).

Distance of epicenter from site: 380 miles

November 5, 1886: Another shock centered at Charleston, felt over the same area as second shock of October 22 (2).

Distance of epicenter from site: 380 miles

June 20, 1893: At 10:07 P M., Jacksonville experienced a slight shock lasting about 10 seconds (1).

Distance of epicenter from site: 230 miles Epicentral Intensity: iv MM(l)

October 31, 1900: Jacksonville, Florida, N. Latitude 30.4, W. Longitude 81.7.

Felt with intensity V MM. Eight distinct shocks felt. No damage (2)

Distance of epicenter from site: 230 miles

Epicentral Intensity: V MM (2)

January 23, 1903: Felt at Tybee Island, Savannah, Georgia N. Latitude 32.1, W. Longitude 81.1, with intensity VI MM; Charleston, S. C.,

IV-V MM; Columbia and Augusta, Ga., III-IV MM. Houses strongly shaken (2)

Distance of epicenter from site: 320 miles

Epicentral Intensity: VI MM (2)

June 12, 1912: This shock centered at Summerville, S. C., N. Latitude 32.9, W. Longitude 80.0, with intensity VII MM and was felt at Wilmington, N. C. Chimneys shaken down at Summerville (2)

Distance of epicenter from site: 380 miles

Epicentral Intensity: VII MM (2)

June 20, 1912: A shock felt strongly at Savannah, Georgia, N. Latitude 32.0, W. Longitude 81 (2).

Distance of epicenter from site: 320 miles

Epicentral Intensity: V MM (2)

1930: A shock felt in Everglades, La Belle, and Ft. Myers. Its seismic origin has been questioned and blasting given as the cause. There is also a report of

windows and dishes rattling at Marco Island about this same time. The probable

intensity at Marco was V MM (l).

Distance of epicenter from site: 130 miles estimated

2.5-87 TABLE 2.5-5 (Cont'd)

November 13, 1935: Two short tremors were felt at Palatka the first at 10:10 P.M., and the second lasting 15 seconds, at 10:30. The second shock

was felt at St. Augustine and on nearby Anastasia Island, but apparently the

disturbance did not extend far as other cities on the coast reported that they

had no disturbances (l).

Distance of epicenter from site: 185 miles

Epicentral Intensity: IV MM - V MM (l)

January 19, 1942: Several shocks occurring near Lake Okeechobee.

Tremors also felt at Miami, Everglades and Ft. Myers (l).

Distance of epicenter from site: 100 miles estimated

Epicentral Intensity: IV MM estimated

November 27, 1973: A shock centered at N. Latitude 28.7, W. Longitude 81.0 (4).

Distance of epicenter from site: 115 miles

Epicentral Intensity: V MM (4)

Earthquake epicenter located in peninsular Florida. Epicentral locations for

these events are shown on Figure 2.5-18.

Ref: 1. Campbell, Robert P., "Earthquakes in Florida," Florida Academy of Sciences, Vol. 6, No. 1, 1943. 2. U. S. Coast & Geodetic Survey, "Earthquake History of U.S.," U. S.

Government Printing Office, 1951-1958. 3. Murphy, L.; N.O.A.A. Personal telephone conversation with Law Engineering Testing Company, 1973. 4. Preliminary data from the National Earthquake Information Center (1-28-74).

2.5-88 Amendment No. 25 (04/12)

TABLE 2.5-6

SUMMARY

OF LIQUEFACTION TEST RESULTS (DYNAMIC TRIAXIAL TESTS - STRESS CONTROLLED)

Cycles to Sample Pore Undisturbed Test Cyclic Pressure = Cycles Applied Test Boring or %Passing Test Dry Moisture Test Confining Stress Confining to Back "B" No. No. Depth Recompacted D 50 No 20G Sieve Density Content D r Pressure Ratio Pressure Liquefaction Pressure Coefficient (Ft.) (mm) (pcf) (%) (%) (psf) (psi)

1 B-5 96 U 0.09 15 113.8 33 100 8,000 .468 15 50 42.7 0.98 2 B-19 77-79 R 2.10 6 113.0 16 100 6,800 .573 10 35 33.5 0.95 3 B-19 78 R 2.10 6 113.0 16 100 6,800 .693 6 20 33.5 0.96 4 B-19 114 R 0.09 18 94.7 28 47 8,850 .472 7 15 48.4 0.98 5 B-20 69 R 0.09 12 95.4 27 50 6,160 .546 1 10 29.2 0.96 6 B-20 97 R - - 91.7 30 36 7,960 .464 2 4 42.7 0.98 7 B-20 137 U 0.10 20 104.9 21 83 10,350 .486 20 125 59.6 0.92 8 E-20 137 R 0.10 20 100.3 24 68 10,350 .347 10 40 59.6 1.00 9 B-20 137 R 0.10 20 104.1 22 81 8,590 .347 23 28 59.6 0.99 10 B-20 137 U 0.10 20 104.0 20 65 10,350 .257 45 75 60.0 0.96 11 B-19 134 U 0.13 34 100.6 24 58 10,000 .479 8 15 57.3 0.98 12 B-19 134 U 0.13 34 103.6 21 68 10,000 .443 10 18 57.3 0.98 13 B-19 134 U 0.13 34 97.8 24 49 10,000 .429 8 18 57.3 0.99

14 Composite R - - 98.0 26 55 10,000 .429 4 5 70.0 0.98 15 Composite R - - 98.0 26 55 10,000 .350 6 8 70.0 0.98 16 A-1 12-27 R 0.13 2.5 97.5 22 85 3,500 .286 604 608 17 A-1 12-27 R 0.13 2.5 97.5 22 85 3,500 .543 91 -

18 A-1 12-27 R 0.13 2.5 97.5 22 85 3,500 .685 56 -

19 A-1 12-27 R 0.13 2.5 97.5 22 85 3,500 .927 18 25

Notes: Test Nos. 1 to 15 presented on Figures 2.5-38 and 2.5-39.

Test Nos. 16 to 19 presented on Figures 2.5-41.

2.5-89

Withheld Under 10 CFR 2.390

DELETED FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT UNIT 1 FIGURE 2.5-84 Amendment No. 23 (11/08)

Withheld Under 10 CFR 2.390 Withheld Under 10 CFR 2.390

FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT - UNIT NO. 1 FINAL SAFETY ANALYSIS REPORT APPENDIX 2A BORING LOGS AND DATA SUMMARIES 2A-i Amendment No. 22 (05/07)

Withheld Under 10 CFR 2.390 Withheld Under 10 CFR 2.390

Summary Two sets of atmospheric dispersion factor values associated with an over-water dispersion model have been provided. One set of values represents a straight line distance calculation from the St. Lucie plant directly to the nearest west bank of the Indian River (Trajectory A). The other set of values represents a circuitous distance calculation from the St. Lucie plant to the nearest shoreline, a meandering trajectory over Big Mud Creek and then a straight line distance to the nearest west bank of the Indian River (Trajectory B). For all the initial overland distances, the atmospheric dispersion characteristics are based on the standard conservative ground release equation. Two assumptions regarding the over-water modification distance have been incorporated in the over-water dispersion model namely, an immediate transition to over-water dispersion characteristics at the shoreline and a critical distance transition prior to over-water modification. The resulting atmospheric dispersion factor values (sec/m3) are: Direct Distance Shoreline Transition Over-Water Transition Trajectory A 1.67 miles 1.32 x 10-4 1.20 x 10-4 Trajectory B 1.77 miles 1.47 x 10-4 1.30 x 10-4 The over-water transition /Qs are representative of what would be expected, whereas, the shore line transition values should be considered a bounding upper limit calculation. It should also be noted that the method utilized for determination of the critical distance is considered conservative. Larger values are likely to be realized. 2H-11 Amendment No. 24 (06/10)

APPENDIX 2I SHORT-TERM (ACCIDENT) ATMOSPHERIC DISPERSION FACTORS FOR THE EXCLUSION AREA BOUNDARY AND LOW POPULATION ZONE FOR AST 2I-i Amendment No. 24 (06/10)

Short-Term (Accident) Atmospheric Dispersion Factors for the Exclusion Area Boundary and Low Population Zone Objective Conservative values of atmospheric dispersion factors at the exclusion area boundary (EAB) and the low population zone (LPZ) were calculated for appropriate time periods using meteorological data collected onsite during the time period including the years 1997, 1998, 1999, 2002 and 2003 which were deemed to represent the most reliable data sets. Methodology The methodology used for this calculation is consistent with Regulatory Guide 1.145 as implemented by the PAVAN computer code (Reference 2). Using joint frequency distributions of wind direction and wind speed by atmospheric stability, the PAVAN computer code provides relative air concentration (/Q) values as functions of direction for various time periods at the EAB and LPZ. Three procedures for calculation of /Qs are utilized for the site boundary and LPZ; a direction-dependent approach, a direction-independent approach, and an overall site /Q approach. The /Q calculations are based on the theory that material released to the atmosphere will be normally distributed (Gaussian) about the plume centerline. A straight-line trajectory is assumed between the point of release and all distances for which /Q values are calculated. The theory and implementing equations employed by the PAVAN computer code are documented in Reference 2. Calculations PAVAN Computer Code Input Data The minimum EAB distance assumed for all directions is 0.97 miles from the center of the Unit 1 containment building. The LPZ distance is taken as 1 mile from the center of the Unit 1 containment building in all directions. All of the releases were considered ground level releases because the highest possible release elevation is from the plant stack at 184 ft. From Section 1.3.2 of Reference 1, a release is only considered a stack release if the release point is at a level higher than two and one-half times the height of adjacent solid structures. For the St. Lucie plant, the elevation of the top of the Unit 1 containment is 225.5 ft. Therefore, the highest possible release point is not 2.5 times higher than the adjacent containment buildings, and thus all releases were considered ground level releases. As such, the release height was set equal to 10.0 meters as required by Table 3.1 of Reference 2. The building cross-sectional area used for the building wake term was 1,565 m2. This area was calculated to be conservatively small in that the height used in the area calculation was from the highest roof elevation of a nearby building to the elevation of the bottom of the containment dome. 2I-1 Amendment No. 26 (11/13)

The tower height at which the wind speeds were measured is 10 m and 57.9 m above plant grade. The wind speed units are given in miles per hour, therefore, the PAVAN variable UCOR was set equal to 101 to convert the wind speeds to meters per second as described in Table 3.1 of Reference 2. The maximum wind speed in each wind speed category was chosen to match the raw joint frequency distribution data, which conforms to the wind speed bins in Table 1 of Reference 3. The maximum wind speed values are 1, 3, 7, 12, 18, 24, and 30 mph. The maximum windspeed in each windspeed category was chosen to match the recommendation of RIS-2006-4. (Reference 4) Results PAVAN computer runs for the EAB and LPZ boundary distances were performed using the data discussed previously. Per Section 4 of Reference 1, the maximum /Q for each distance was determined and compared to the 5% overall site value for the boundary under consideration. The maximum EAB and LPZ /Qs that resulted from this comparison are provided in the table below: Exclusion Area Boundary and Low Population Zone /Qs Time Period EAB /Q (sec/m3) LPZ /Q (sec/m3) 0-2 hours 9.84E-05 9.56E-05 0-8 hours 5.53E-05 5.34E-05 8-24 hours 4.15E-05 3.99E-05 1-4 days 2.22E-05 2.12E-05 4-30 days 9.06E-06 8.55E-06 References 1. USNRC Regulatory Guide 1.145, "Atmospheric Dispersion Models for Potential Accident Consequence Assessments at Nuclear Power Plants," Revision 1, November 1982. (Reissued February 1983 to correct page 1.145-7). 2. NUREG/CR-2858, "PAVAN: An Atmospheric Dispersion Program for Evaluating Design Basis Accidental Releases of Radioactive Materials for Nuclear Power Stations," November 1982. 3. Safety Guide 23, "Onsite Meteorological Programs," February 17, 1972. 4. USNRC Regulatory Issue Summary RIS-2006-4, "Experience with Implementation of Alternative Source Terms," March 7, 2006. 2I-2 Amendment No. 26 (11/13)

APPENDIX 2J SHORT-TERM (ACCIDENT) ATMOSPHERIC DISPERSION FACTORS FOR THE CONTROL ROOM FOR AST 2J-i Amendment No. 24 (06/10)

Short-Term (Accident) Atmospheric Dispersion Factors for the Control Room Objective Conservative values of atmospheric dispersion factors for the control room were calculated for appropriate time periods using meteorological data collected onsite during the time period including the years 1997, 1998, 1999, 2002 and 2003 which were deemed to represent the most reliable data sets. Methodology The ARCON96 computer code is used by the USNRC staff to review licensee submittals relating to control room habitability (Reference 1). Therefore, the ARCON96 computer code was used to determine the relative concentrations (/Qs) for the control room air intakes and inleakage locations. The ARCON96 computer code uses hourly meteorological data for estimating dispersion in the vicinity of buildings to calculate relative concentrations at control room air intakes that would be exceeded no more than five percent of the time. These concentrations are calculated for averaging periods ranging from one hour to 30 days in duration. The theory and implementing equations employed by the ACRCON96 computer code are documented in Reference 1. Calculations/ARCON Computer Code Input Data Five years of meteorological data were used for the ARCON96 computer code runs. A number of various release-receptor combinations were considered for the control room /Qs. These different cases were considered to determine the limiting release-receptor combinations for the various events. The case matrix for these combinations is provided in Table 2J-2. The distance and direction inputs for the ARCON96 runs may be found in Table 2J-1. The distances were converted from feet to meters with a factor of 0.3048 m/ft. The distances in meters were then rounded down to the nearest tenth for conservatism. The elevation difference term was set equal to zero for each case since all elevation points are taken with respect to the same datum. The intake heights were determined as the intake elevations less the plant grade elevation of 19 ft. The lower and upper measurement heights for the meteorological data were entered as 10 m and 57.9 m, respectively, for each case. The mph option was selected for the wind speed units. 2J-1 Amendment No. 26 (11/13)

A ground level release was chosen for each scenario since none of the release points are 2.5 times taller than the closet solid structure as called out in Section 3.2.2 of Reference 3 for stack releases. The top of the containment structures is at an elevation of 225.5 ft. The highest release point is from the top of the plant stack at an elevation of 184 ft., which is not 2.5 times higher than the nearby containment structure. The vertical velocity, stack flow, and stack radius terms were all set equal to zero since each case is a ground level release. The vent release option was not selected for any of the scenarios. The actual release height was used in the cases. No credit was taken for effective release height due to plume rise; therefore, for the releases from the stacks, the release elevations were set equal to the stack top elevation. The release heights were taken as the release elevations less the plant grade elevation of 19 ft. The only cases in this analysis that take credit for the building wake effect are the scenarios where the release is from the containment building. Some of the other scenarios have buildings between the release and receptor points, but for these cases the building wake was not credited for the sake of conservatism. Not crediting wakes was accomplished by setting the building area term equal to 0.01 m2 as stated in Table A-2 of Reference 3. The building area used is a conservatively determined containment cross sectional area. The width used is equal to the inside diameter of the containment building plus the thickness of the wall, while the height is taken as the distance between the top of the cylinder potion of the containment structure and the highest auxiliary building roof elevation. This building cross-sectional area is equal to 1,565 m2. All of the default values in the ARCON96 code were unchanged from the code default values with the following exceptions as recommended in Table A-2 of Reference 3: A value of 0.2 is used for the surface roughness length, m, in lieu of the default value of 0.1, and A value of 4.3 is used for the averaging sector width constant, in lieu of the default value of 4.0. The minimum wind speed was left at 0.5 m/s per the guidance instruction in Table A-2 of Reference 3. Results ARCON96 computer runs for the various release points and control room intake locations were performed using the data discussed previously. Per Reference 3, the 95th percentile /Q values were determined. The resulting /Qs are listed in Table 2J-2. References 1. NUREG/CR-6331 PNL-10521, "Atmospheric Relative Concentrations in Building Wakes," May 1995, with Errata dated July 1997. 2. Safety Guide 23, "Onsite Meteorological Programs," February 17, 1972. 3. USNRC Regulatory Guide 1.194, "Atmospheric Relative Concentrations for Control Room Radiological Habitability Assessments at Nuclear Power Plants," June 2003. 2J-2 Amendment No. 24 (06/10)

TABLE 2J-1 Direction and Distance Data Release Point Receptor Point Release Height (ft) Release Height (m) Receptor Height (ft) Receptor Height (m) Distance (ft) Distance (m) Direction With Respect to True North Stack/Plant Vent N CR intake 184 56.1 59.75 18.2 48.08 14.6 58 Stack/Plant Vent S CR intake 184 56.1 59.75 18.2 126.69 38.6 354 RWT N CR intake 48.22 14.6 59.75 18.2 245.31 74.7 65 RWT S CR intake 48.22 14.6 59.75 18.2 263.64 80.3 39 FHB Closest Point N CR intake 43.25 13.2 59.75 18.2 120.6 36.7 48 FHB Closest Point S CR intake 43.25 13.2 59.75 18.2 184.26 56.1 11 Aux. Bldg. Louver L-7B N CR intake 38.17 11.6 59.75 18.2 123.77 37.7 72 Aux. Bldg. Louver L-7A S CR intake 38.17 11.6 59.75 18.2 136.97 41.7 34 Condenser N CR intake 5.25 1.6 59.75 18.2 153.24 46.7 245 Closest ADV N CR intake 53 16.1 59.75 18.2 105.68 32.2 306 Closest ADV S CR intake 53 16.1 59.75 18.2 214.82 65.4 319 Closest Feedwater Line Point N CR intake 17 5.2 59.75 18.2 83.29 25.3 306 Closest Feedwater Line Point S CR intake 17 5.2 59.75 18.2 193.15 58.8 321 2J-3 Amendment No. 24 (06/10)

TABLE 2J-1 Direction and Distance Data Release Point Receptor Point Release Height (ft) Release Height (m) Receptor Height (ft) Receptor Height (m) Distance (ft) Distance (m) Direction With Respect to True North Containment Maintenance Hatch N CR Intake 16 4.9 59.75 18.2 172.4 52.5 359 Containment Maintenance Hatch S CR Intake 16 4.9 59.75 18.2 279.09 85.0 348 FHB Closest Point Midpoint Between Intakes 43.25 13.2 59.75 18.2 142.19 43.3 25 Stack/Plant Vent Midpoint Between Intakes 184 56.1 59.75 18.2 74.85 22.8 8 RWT Midpoint Between Intakes 48.22 14.6 59.75 18.2 244.91 74.6 52 Aux. Bldg. Louver L-7A Midpoint Between Intakes 38.17 11.6 59.75 18.2 118.59 36.1 59 Closest ADV Midpoint Between Intakes 53 16.1 59.75 18.2 160.26 48.8 314 Closest Feedwater Line Point Midpoint Between Intakes 17 5.2 59.75 18.2 138.15 42.1 315 Containment Maintenance Hatch Midpoint Between Intakes 16 4.9 59.75 18.2 223.66 68.1 351 2J-4 Amendment No.24 (06/10)

TABLE 2J-1 Direction and Distance Data Notes: 1. Release heights are calculated as 19 feet less than the reference elevations to account for the plant grade elevation. 2. The FHB closest point release elevation is taken as the roof elevation since the SW corner of the roof is the closest building point to the intakes. 3. Release and receptor points are considered to be at the centerpoint or centerline of all openings. 4. The only release/receptor combination that does not have the intakes in the same wind direction window from the release point is for the releases from the plant stack. All other release points analyzed result in both control room intakes being in the same wind direction window. Therefore, credit may be taken for intake dilution only for releases from the plant stack. 5. The receptor point for the "midpoint between intakes" is taken as being on the outside of the control room (and H&V room) east wall. The receptor elevation is taken as the average of the receptor elevations for the two outside air intakes. 6. Atmospheric dispersion factors for the releases to the midpoint between the control room intakes are required for the limiting cases to be used during the time period when the control room intakes are isolated. This midpoint receptor location is used to calculate the /Q value to be used for the unfiltered control room inleakage dose. 7. The closest containment/shield building penetration to the intakes that is directly exposed to the atmosphere is the closest feedwater line penetration. 2J-5 Amendment No. 24 (06/10)

TABLE 2J-2 Control Room /Qs This table summarizes the results for /Q factors for the control room intakes for the various accident scenarios. Values are presented for the unfavorable intake prior to intake isolation, the midpoint between the intakes for during isolation, as well as values for the favorable intake due to the manual selection of the favorable control room intake after unisolation and initiation of filtered air make-up. These values are not corrected for Control Room Occupancy Factors but do include taking credit for dilution where allowed. Section 3.3.2.3 of Reg. guide 1.194 provides the following guidance for dual intake ventilation systems which allow the operator to manually select the least contaminated outside air intake as a source of outside air makeup and close the other intake: "The /Q value for the limiting intake should be used for the time interval prior to intake isolation. This /Q value may be reduced by a factor of 2 to account for dilution by the flow from the other intake. The /Q values for the favorable intake are used for the subsequent time intervals. The /Q values may be reduced by a factor of 4 to account for the duel inlet and the expectation that the operator will make the proper intake selection. This protocol should be used only if the dual intakes are in different wind direction windows." Based on the layout of the site, the only cases that may take credit for dilution are when the releases are from the plant vent stack. Therefore, the Plant Stack values shown below include a reduction by a factor of 2 prior to isolation and a reduction by a factor of 4 after the control room is aligned to the favorable intake for filtered make-up. However, dilution is not credited during the time period when the control room intakes are isolated for these cases. Indicates credit for dilution taken for this case. The atmospheric dispersion factors corresponding to ADVs were determined to be more limiting than those from the MSSVs for all time periods. Therefore, the more limiting ADV values have been used throughout the analyses for all secondary releases. No distinction is made between automatic steam relief from the MSSVs and controlled releases from the ADVs for radiological purposes. Release- Receptor Pair Release Point Receptor Point 0-2 hour /Q (sec/m3) 2-8 hour /Q (sec/m3) 8-24 hour /Q (sec/m3) 1-4 days /Q (sec/m3) 4-30 days /Q (sec/m3) A Stack/Plant Vent N CR Intake 2.39E-03 B Stack/Plant Vent S CR Intake 6.93E-04 4.88E-04 2.19E-04 1.46E-04 1.28E-04 C RWT N CR Intake 1.37E-03 D RWT S CR Intake 1.12E-03 9.10E-04 3.84E-04 2.93E-04 2.37E-04 E FHB Closest Point N CR Intake 4.99E-03 F FHB Closest Point S CR Intake 2.01E-03 1.44E-03 6.25E-04 4.34E-04 3.33E-04 G Aux. Bldg. Louver L-7B N CR Intake 4.80E-03 H Aux. Bldg. Louver L-7A S CR Intake 3.61E-03 2.87E-03 1.20E-03 9.07E-04 7.13E-04 I Condenser SJAE N CR Intake 3.02E-03 J Closest ADV N CR Intake 6.30E-03 2J-6 Amendment No. 26 (11/13)

TABLE 2J-2 Release- Receptor Pair Release Point Receptor Point 0-2 hour /Q (sec/m3) 2-8 hour /Q (sec/m3) 8-24 hour /Q (sec/m3) 1-4 days /Q (sec/m3) 4-30 days /Q (sec/m3) K Closest ADV S CR Intake 1.62E-03 1.32E-03 5.06E-04 3.88E-04 3.30E-04 L Closest Feedwater Line Point N CR Intake 7.29E-03 M Closest Feedwater Line Point S CR Intake 1.76E-03 1.41E-03 5.72E-04 4.29E-04 3.57E-04 N Stack/Plant Vent Midpoint Between Intakes 3.91E-03 O RWT Midpoint Between Intakes 1.34E-03 P Aux. Bldg. Louver L-7A Midpoint Between Intakes 5.03E-03 Q Closest ADV Midpoint Between Intakes 2.84E-03 R Closest Feedwater Line Point Midpoint Between Intakes 3.17E-03 S Containment Maintenance Hatch N CR Intake 1.90E-03 T Containment Maintenance Hatch S CR Intake 8.22E-04 6.57E-04 2.87E-04 1.92E-04 1.74E-04 U Containment Maintenance Hatch Midpoint Between Intakes 1.21E-03 V FHB Closest Point Midpoint Between Intakes 3.27E-03 2J-7 Amendment No. 26 (11/13)

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