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Enclosuflooding Hazard Reevaluation Report, Cover Through Page 2.3-1 Through End
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Site:	South Texas
Issue date:	03/11/2013
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Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 2.3 Dam Breaches and Failures For flooding hazards associated with dam and embankment failures, the UFSAR of STP 1 &2 (Reference 2.3-1) identifies two sources that are most critical to the plant. They are: (1) the postulated cascade failure of the major upstream dams on the Colorado River; and (2) the breaching of the embankment of the onsite Main Cooling Reservoir (MCR).

These two sources also form the primary basis of the flooding reevaluation on dam breaches and failures for STP 1 & 2 because there are no new dams or large water impoundments (including the previously proposed Columbus Bend Dam) planned for the Colorado River in the next 45 to 50 years, according to the 2007 State Water Plan (Reference 2.3-17), that could produce a more severe flooding risk at the site.

The reevaluation of upstream dam failure scenarios and the postulated flood risk to the STP 1 &

2 site are described in Subsection 2.3.1. Postulated flood risk from the MCR failure is described in Subsection 2.3.2.

The Essential Cooling Pond (ECP), a man-made excavated pond 9 ft deep with an 8-foot-high embankment completely surrounding its perimeter as described in Section 9.2 of the UFSAR (Reference 2.3-1), is not considered a realistic or critical source of flooding that could impact the safety functions of the plant. The ECP is the ultimate heat sink sized to have a 30-day water supply for the Essential Cooling Water System (ECWS) to support the safe shutdown of both units. Water circulation within the pond flows clockwise and is controlled by a central dividing dike (Reference 2.3-1, Subsection 2.5.6). The crest elevations of the circumferential embankment and the dividing dike are at El. 34 ft MSL and El. 38 ft MSL, respectively. Both the dividing dike and the southern embankment are seismic Category I in addition to being designed to withstand the effects of a breach in the MCR embankment. The remaining portion of the northern embankment is designed to withstand the effects of the Colorado River dam failures. Failures of the ECP embankments are therefore not expected. However, in the highly unlikely event of an ECP embankment breach, the flooding hazard to the plant would have a much smaller impact and be bounded by the hypothetical breaching of the MCR embankment evaluated in Subsection 2.3.2 because:

(a) The normal operating elevation of the ECP is between El. 25.6 ft MSL and El. 26.0 ft MSL (UFSAR Subsection 9.2.5), about 2 ft lower than the plant grade of El. 28 ft MSL and much lower (by about 23 ft) than the maximum operating level of El. 49 ft MSL for the MCR (UFSAR Subsection 2.4.8);

(b) In the event of a severe antecedent storm event, the starting ECP water level prior to the hypothetical embankment breach would still be significantly lower than that of the MCR; (c) The ECP contains a surface area of 46.5 acres and a storage volume of approximately 112,000,000 gallons (344 acre-ft) at El. 25.5 ft MSL (UFSAR Subsection 9.2.5), which are multiple orders of magnitude smaller than the nominal 7,000 acres of surface area and 202,700 acre-ft of storage volume of the MCR at its normal maximum operating level of El. 49 ft MSL (UFSAR Figure 2.4.8-7); and (d) The topography at the site slopes gently from north to south, which would divert the flood wave from an ECP embankment breach away from the power block, especially with the natural ground elevation near the ECP at approximately El. 25 ft MSL (UFSAR Dam Breaches and Failures 2.3-1

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STPJ & 2 Fukushimna Response Project Subsection 2.5.6) and finish grade at about El. 27 ft MSL, lower than the plant grade of El. 28 ft MSL at the power block (UFSAR Subsection 2.4.4).

Based on the assessment of the above factors, it can therefore be concluded that an ECP embankment breach will have no flooding impact on the safety functions of STP 1 & 2.

2.3.1 Upstream Dam Failures The potential flooding hazards on the STP site due to upstream dam failures have been evaluated in UFSAR of STP 1 &2 (Reference 2.3-1) and also in a more recent (2007) analysis performed to support Combined License Application (COLA) of the proposed STP 3 & 4 (Reference 2.3-2).

The flooding reevaluation for STP 1 & 2 adopts the approach and methodology for estimating the flood wave generated by upstream dam failures in the STP 3 &4 COLA analysis (Reference 2.3-2), which provides a comprehensive and conservative assessment of the upstream dam failure scenario using the most current data available and the industry standard numerical modeling tool, HEC-RAS of the U.S. Army Corps of Engineers (USACE) (Reference 2.3-9).

Review of the COLA analysis indicates that it still provides a bounding analysis and meets the objectives of the flooding reevaluation of using present-day methodology and data. In particular, the STP 3 &4 COLA analysis follows the guidance of ANSI/ANS 2.8-1992 (Reference 2.3-7) on the assessment of the potential dam failure modes and the specification of the antecedent and combined event conditions, consistent with the recommendations of NUREG/CR-7046. Since the recent completion of the COLA analysis in 2007, no major upstream dam or impoundment on the Colorado River or tributaries has been constructed or proposed, and no change to the hydrologic and hydraulic properties of the dams and reservoirs evaluated, as well as those of the affected channel and floodplain, is identified or expected. Further, there is no new hydraulic control or modifications on the Lower Colorado River that could potentially affect the boundary conditions used in the HEC-RAS model simulation.

The combined effects of wind setup and wave runup from the STP 3 & 4 COLA analysis, however, are not directly applicable to the STP 1 & 2 facilities because of the differences in site specific properties such as local grade elevations and slopes, surface roughness and fetches.

Therefore, wind setup and wave runup are reanalyzed specifically for STP 1 & 2 safety facilities, including the Essential Cooling Water (ECW) intake structure, for a two-year design wind speed occurring coincidently with the maximum "still" water level as a result of the postulated upstream dam break scenario, per ANSI/ANS-2.8-1992 (Reference 2.3-7).

The STP I & 2 site is located on the west bank of the Colorado River in Matagorda County, Texas, about 10.5 river miles upstream of the Gulf Intracoastal Waterway (GIWW). There are a total of 68 dams with storage capacity in excess of 5000 acre-feet (AF) on the Colorado River and its tributaries upstream of the STP site. These dams and reservoirs are owned and operated by different entities including the Lower Colorado River Authority (LCRA), the U.S.

Bureau of Reclamation (USBR), the Colorado River Municipal Water District (CRMWD), other local municipalities and utilities. Figures 2.3-1a and 2.3-1b show the locations of the 68 dams.

Specific information of these dams that are relevant to the flood risk assessment of STP 1 & 2 are summarized in Table 2.3-1, based on data collected primarily from the Texas Water Development Board (TWDB), Texas Commission for Environmental Quality (TCEQ), and LCRA.

The six hydroelectric dams - Buchanan, Roy Inks, Alvin Wirtz, Max Starcke, Mansfield, and Tom Miller, owned and operated by LCRA are known as the Highland Lake dams.

Damn Breaches and Failures 2.3-2

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STP1 & 2 Fukushitna Response Project In Texas, both private and public dams are monitored and regulated by TCEQ under the Dam Safety Program. Existing dams, as defined in Rule §299.1 Title 30 of the Texas Administrative Code (Reference 2.3-3), are subject to periodic re-evaluation in consideration of continuing downstream development. Rule §299.14 of Title 30 (Table 3) on Hydrologic Criteria for dams stipulates the minimum acceptable spillway evaluation flood (SEF) for re-evaluating dam and spillway capacity for existing dams to determine whether upgrading is required. Similarly, on the structural considerations, evaluation of an existing dam includes, but is not limited to, visual inspections and evaluations of potential problems such as seepage, cracks, slides, conduit and control malfunctions, and other structural and maintenance deficiencies which could lead to failure of a structure.

Following the 1987 National Dam Safety Inspection Program recommendations of the Texas Water Commission, a predecessor agency of the TCEQ, to upgrade two of the Highland Lake dams due to unsafe condition, LCRA initiated a program to evaluate all six Highland Lake dams with respect to hydrologic, structural and geotechnical criteria.

In 1990, LCRA began a 15-year plan of Dam Modernization Program to address the safety condition of five of the six dams. A 1992 dam safety evaluation study commissioned by LCRA (Reference 2.3-4) indicates that Wirtz, Starcke, and Tom Miller Dams would be overtopped during a Probable Maximum Flood (PMF) event, and certain sections of Buchanan, Wirtz, and Tom Miller Dams could have instability problems during severe flood conditions. The concrete dam sections of Mansfield Dam, however, would be stable during the PMF. At the completion of LCRA's Dam Modernization Program in January of 2005, substantial upgrade work had been undertaken at Buchanan, Inks, Wirtz, and Tom Miller Dams to address the unsafe conditions (Reference 2.3-5). Upgrade at Mansfield Dam was considered not necessary as it is able to withstand the PMF without further reinforcement. Even in the event of failures of either Buchanan, Inks, Wirtz, or Starcke dams, Mansfield Dam would hold their flood volumes without overtopping (Reference 2.3-6).

2.3.1.1 Dam Failure Permutations 2.3.1.1.1 Failures of Upstream Dams on the Colorado River Of all the dams on the Colorado River upstream of the STP 1 & 2 site, Mansfield Dam would generate the most significant dam break flood risk on the site. Mansfield Dam has the largest dam height of 266.4 ft and the largest reservoir storage capacity of 3.3 million acre-feet (MAF),

at top of the dam. Among all the dams upstream, Mansfield Dam is also closest to the site at about 305 river miles upstream of the STP 1 & 2 site. The next major dam upstream that could pose significant flood risk to the site is the Buchanan Dam located at about 402 river miles upstream of STP 1 & 2. It has a height of 145.5 ft and a top-of-dam storage capacity of 1.18 MAF. Further upstream, the Simon Freese Dam, with a height of 148 ft and a top-of-dam storage capacity of 1.47 MAF, and the Twin Buttes Dam, with a height of 134 ft and top-of-dam storage capacity of 1.29 MAF are considered to have major, though not as significant, contribution to the flood risk at the STP site. They are located at about 199 miles and 290 miles, respectively, upstream of Buchanan Dam.

There are two failure permutations postulated of the upstream dams:

Dam Breaches and Failures 2.3-3

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STPI & 2 Fukushimna Response Project Scenario No. 1 - Simultaneous failure of all upstream dams induced by a seismic event. The failure is to occur coincidentally with a 2-year design wind event and a 500-year flood or a one-half probable maximum flood (PMF) per American National Standard ANSI/ANS-2.8-1992 (Reference 2.3-7).

Scenario No. 2 - Domino-type failure of upstream dams with the same coincidental wind and flood events as in Scenario No. 1. It is postulated that the upstream-most dam(s) would fail first, thereby releasing a dam break flood wave (or waves) that propagates downstream and triggers the failure of the downstream dams one after another in a cascading manner. It is assumed that the 56 dams on the Colorado River and its tributaries upstream of Buchanan Dam (with top-of-dam capacity over 5000 AF) would fail in such a manner that their flood flow, expressed in terms of their respective top-of-dam storage volumes, would arrive at Lake Buchanan at approximately the same time, triggering the failure of Buchanan Dam. The dam break flood flow from Buchanan Dam would then propagate downstream to Lake Travis, overtopping Mansfield Dam and causing it to fail. The dam break flood from Mansfield Dam then propagates downstream to the STP 1 & 2 site. The failure is to occur coincidentally with a 2-year design wind event and a 500-year flood or a one-half probable maximum flood (PMF) per American National Standard ANSI/ANS-2.8-1992 (Reference 2.3-7).

Three upstream dams, Inks, Wirtz, and Starcke, located between Buchanan and Mansfield Dams, and two other upstream dams, Tom Miller and Longhorn Dams, located at 20 miles and 27 miles downstream of Mansfield Dam, were not included in the dam break analysis as their dam heights and potential flood volumes would have insignificant impact on the flood risk as compared to Mansfield Dam or Buchanan Dam.

There are five "off-channel" dams located on the tributaries of the Colorado River between Mansfield Dam and the STP site. They are: Decker Creek Dam (Lake Long), Bastrop Dam, Cummins Creek WS SCS Site 1 Dam, Cedar Creek Dam (Fayette Reservoir), and Eagle Lake Dam. These off-channel storage dams were also assumed to have no effect on the maximum dam break flood level at the STP 1 &2 site, as compared to the major dams on the main stem of the Colorado River.

Of these two permutations, Scenario No. 2 would generate the most critical, and therefore bounding, flood level at STP 1 & 2 because of the deliberate alignment of the travel and arrival of the dam breach flood volumes and flood peaks from the major upstream dams.

Consequently, only the flood risk resulting from Scenario No. 2 was further evaluated. Upstream dam failures induced by hydrologic causes such as probable maximum flood (PMF) will not be the controlling scenario in the evaluation of the maximum flood risk at the STP site. This is because the large dams with high hazard potential, such as O.C. Fischer, Simon Freese, Buchanan and Mansfield Dams, as listed in Table 2.3-1, were either designed or have been upgraded to accommodate and sustain their respective PMFs in accordance with the hydrologic criteria for dams as defined in Rule 299.14 Title 30 of the Texas Administrative Code (Reference 2.3-3). Mansfield Dam, in particular, would be able to hold the dam break flood volumes of either Buchanan, Wirtz, or Starcke dams. Besides, the assumption that a domino-type dam failure of the 56 dams upstream of Buchanan with an aggregated top-of-dam storage volume of 6.87 MAF all arriving at Buchanan at about the same time is highly conservative and would have bounded the potential flood risk caused by hydrological dam failures.

Dam Breaches and Failures 2.3-4

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project 2.3.1.2 Unsteady Flow Analysis of Potential Upstream Dam Failures 2.3.1.2.1 Colorado River Dams Table 2.3-1 lists the height, length, top-of-dam storage capacity, type, and year of completion of the 68 dams with a top-of-dam storage capacity larger than 5000 AF each. Of these 68 dams, Mansfield Dam, Buchanan Dam and 56 other dams upstream of Buchanan Dam were selected for inclusion in the dam break analysis. Dams with less than 5000 AF storage capacity, i.e., less than 0.2% of that of Mansfield Dam, were excluded from further evaluation as the impact of their potential breaching on the flood risk at the site would be minimal. The top-of-dam storage volume of Mansfield Dam is about 3.3 MAF, estimated from the elevation-storage capacity curves given in Reference 2.3-8. Similarly, the top-of-dam storage volume of Buchanan Dam is estimated to be about 1.18 MAF. The combined top-of-dam-storage volume of the 56 dams upstream of Buchanan Dam is 6.87 MAF.

2.3.1.2.1.1 Conceptual Unsteady Flow Analytical Model The dam breach option of the USACE River Analysis System computer program (HEC-RAS)

(Reference 2.3-9) was used to simulate the dam breach flood waves, which were then routed downstream to the STP 1 & 2, using the unsteady flow option of the program.

In the conceptual dam break flood model, the 56 dams upstream of Buchanan Dam would fail in a domino manner, with their combined top-of-dam storage capacity, totaling 6.87 MAF, arriving at Buchanan Dam at approximately the same time. As the flood level at Buchanan Dam rises to about 3 ft over the dam crest elevation of 1025.35 ft MSL, the dam would fail, thereby releasing the flood storage of Buchanan Dam plus the combined flood volumes from the 56 upstream dams. In accordance with the combined events requirements stipulated in the American National Standard ANSI/ANS-2.8-1992 (Reference 2.3-7), the evaluation of potential flood risks as a result of non-hydrologic dam break failures should also consider a coincidental event equal to a 500-year flood or one-half probable maximum flood (PMF), whichever is less. In this analysis, a constant flood flow of 500,000 cfs, slightly higher than the peak Standard Project Flood (SPF) inflow at Buchanan Dam and the 500-year flood peak inflow at Mansfield Dam, was conservatively used to represent the coincidental flow. The SPF and 500-year flood flow at several locations on the Colorado River are listed in Table 2.3-2. They were estimated by Halff Associates, Inc. as part of the Lower Colorado River flood damage evaluation project conducted for LCRA and Fort Worth District Army Corps of Engineer (Reference 2.3-10). The 500,000 cfs coincidental flow was applied to the entire model reach from Buchanan Dam to the downstream boundary at 4600 ft (0.9 river miles) upstream of the Gulf Intracoastal Waterway.

The flood wave from the breaching of Buchanan Dam would propagate down to the 266.4-ft high Mansfield Dam, with a crest elevation at 754.1 ft MSL and a top-of-dam storage capacity of 3.30 MAF. (In 1941, a 4-ft parapet wall was added to the dam crest raising its elevation from 750.1 ft MSL to 754.1 ft MSL to provide additional flood storage capacity.) Mansfield Dam was postulated to fail when it was overtopped by 3 ft at El. 757.1 ft MSL. The three dams located between Buchanan and Mansfield Dams: Roy Inks, Alvin Wirtz, and Max Starcke dams, have a combined storage of about 298,300 AF. These dams were not assumed to fail in the dam break model because their combined total storage amounts to only about 9% of the total dam break flood volume at Mansfield. The SPF flood hydrographs from 19 tributaries between Buchanan and Mansfield dams as estimated by Halff Associates, Inc. in the flood damage evaluation study (Reference 2.3-10) were included as tributary inflows to this reach. The tributary inflows Dam Breaches and Failures 2.3-5

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STPI & 2 Fukushima Response Project together with the dam break flood wave from Mansfield Dam were then routed to the STP 1 & 2 site in the HEC-RAS model.

2.3.1.2.1.2 Physical Dam Data and Estimates of Breached Sections Buchanan Dam, located at about 402 river miles upstream of STP 1 & 2, is 10,987 ft in length. It has two separate multiple concrete arch sections as well as a number of gravity sections (Reference 2.3-8). The main dam section consists of 29 concrete arches, each of 70 ft in width and 145.5 ft in height. The total length of this multiple concrete arch section is 2030 ft and it occupies the deepest part of the river channel. To the right (looking downstream) is another shorter multiple concrete arch section of 805 ft in length, consisting of 23 arches of 35 ft wide each. Following the guidelines from Federal Energy Regulatory Commission (FERC) on dam break analysis (Reference 2.3-11), 15 of the 29 larger arches (70 ft wide each) and 12 of the 23 smaller arches (35 ft wide each) were assumed to breach in the simulation. The breach section in the model was represented by a vertical section with a total width of 1470 ft and extending from the top of the dam to the bottom. The time to complete the breach was assumed to be 0.1 hour1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months , based on the guidelines from FERC for the estimation of the dam breach parameter (Reference 2.3-11). The model cross-section at Buchanan Dam is shown in Figure 2.3-2.

Mansfield Dam, at about 305 river miles upstream of STP 1 & 2, has a 2710 ft long, 266.4 ft high concrete gravity section occupying the main river channel, and a 4380 ft long earthen rockfill saddle section with a maximum height of about 150 ft on the left side (looking downstream) (Reference 2.3-8). The total storage capacity is 3.13 MAF at the dam crest elevation of 750.1 ft MSL. With the installation of the 4-ft parapet wall in 1941, the storage capacity increased to 3.30 MAF. Following the FERC guidelines (Reference 2.3-11), about half of the 2710 ft concrete gravity section was postulated to fail when overtopped by 3 ft, resulting in a 1360 ft wide vertical breached section from top to bottom. The time to complete the breach was also assumed to be 0.1 hour1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months . The model cross-section for Mansfield Dam is shown in Figure 2.3-3.

Table 2.3-3 lists the dam breach characteristics used to model the failure of these two dams.

2.3.1.2.1.3 Channel Geometry The channel geometry in the HEC-RAS dam break model was adopted from the river cross-sectional data of Halff's flood damage evaluation study for the Lower Colorado River (Reference 2.3-10 and discussed in Section 2.2). The Halff model has a total model reach length of 474 river miles represented by 1048 cross-sections from Texas Highway 190 upstream of Buchanan Dam, to a section at 4600 ft (0.9 river miles) upstream of the Gulf Intracoastal Waterway just north of Matagorda Bay. The HEC-RAS dam break model developed for STP 1 & 2 has a shorter river reach of 414 miles starting from Buchanan Dam on the upstream end and was represented by a total of 793 model cross-sections. All bridge crossings specified in the Halff model were removed because they were assumed to be washed away during the dam break event. In addition, all ineffective flow areas as well as levees specified in the Halff model were also removed, when deemed appropriate. The locations of these cross-sections are shown in Figure 2.3-4. The elevations of each of the cross-sections were referenced to the North America Vertical Datum 1988 (NAVD 88) in the Halff study. The HEC-RAS dam break model runs were also conducted in NAVD 88 datum. However, the flood level predictions were converted to MSL (or NGVD 29) for comparison with the STP plant grades.

Dam Breaches and Failures 2.3-6

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP I & 2 Fukushima Response Project Because the top-of-dam storage at Buchanan Dam was estimated to be 1.18 MAF, while the aggregated total top-of-dam storage of the 56 selected dams upstream of Buchanan Dam was estimated to be 6.87 MAF, it would not be possible for Buchanan Dam to accommodate the entire dam break flood volume from the breaching of these upstream dams. In order to properly account for the residual flows that could still arrive at and propagate downstream of Buchanan Dam after its failure, new model cross sections were introduced upstream of Buchanan Dam to extend the model reach by 36 miles to approximate the additional volume required to accommodate the combined dam break flood flow of 6.87 MAF from the dams upstream. The upstream reach extension consists of 37 rectangular cross sections 16,030-ft wide with a bottom elevation at 915.8 ft MSL. The cross-sectional width of 16,030 ft is similar to those of the three cross-sections behind Buchanan Dam in the Halff model (Reference 2.3-10). The total flood volume in the model simulation would be over 8.0 MAF behind Buchanan Dam when it breaches at 3 ft above dam crest.

The primary objectives of the Halff study are for flood damage evaluations of the Lower Colorado River and therefore the model predictions were conducted for flood events up to the SPF. During extreme floods, inter-basin spillage could occur. Flood flow from the Colorado River could overspill into its neighboring sub-basins, such as Tres Palacios River to the west and San Bernard River and Peyton Creek to the east. In the flood of 1913, floodwaters from the Colorado River sub-basin overflowed into Caney Creek sub-basin to the east of the Colorado River near Wharton. With predictably higher flood discharges during the postulated dam failure scenario, the channel cross-sections of the Halff study need to be extended beyond their limits to more accurately reflect the additional floodplain areas that would be inundated during the passage of the dam break flood waves. As HEC-RAS would automatically assume a vertical wall at the pre-set boundaries of the flood channel or floodplain, the extension could mitigate potentially unrealistic flood levels as a result of artificial limitation on the cross-sectional geometries imposed by the model setup. This can have a significant impact on the predicted flood peak in the lower reach of the river near the STP 1 & 2 site, where the drainage divides between sub-basins are relatively low in elevation.

A comparison was made between the simulated water levels from the initial dam break runs and the elevations of the drainage divides to determine the approximate location where inter-basin spillage would occur. It was found that inter-basin spillage could occur near Garwood.

Therefore, about 1.9-mile extension was added to the Halff model cross-sections on each side starting from near Garwood. The width of the extension on each side was gradually increased to about 9.5 miles near Wharton down the river. Because the topography is, in general, higher west of the Colorado River towards the Palacio River sub-basin, the cross-sectional extensions in the downstream reach shifted eastward towards the San Bernard River and the Peyton Creek sub-basins. Eventually, near the STP 1 & 2 site, the river cross-sections were extended towards the east for some 17 miles. Typical model cross-sections at four locations on the model river reach including the extended sections are shown in Figures 2.3-5 to 2.3-8.

The USGS 30-m National Elevation Dataset (NED) digital elevation model data used to establish the cross-sectional extensions was referenced to MSL (or NGVD 1929), while the Halff model was referenced to NAVD 88. As the difference between these two datum references for this reach of the Lower Colorado River is less than 0.3 ft, no corrections to the datum, except for 32 sections, were made to adjust the elevations of the extensions to NAVD 88 datum. The 32 sections with datum corrected were located between the STP site and the downstream boundary and were adopted from the PMF routing model described in Section 2.2.

Dam Breaches and Failures 2.3-7

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project The locations and extents of the cross-sections used in the HEC-RAS dam break model are shown in Figure 2.3-4.

2.3.1.2.1.4 Manning's n Values Used in the HEC-RAS Model The Manning's n values used in the Halff HEC-RAS model were calibrated with historical storms and measured flood levels using the values suggested in Table 2.3-4 (Reference 2.3-10) as initial estimates. The calibrated values are in the range of 0.025 to 0.046 for the river channel and 0.045 to 0.100 for the overbank areas, and they were used in the Halff study to model flood conditions up to the SPF. The extensions in the dam break model adopted the same Manning's n values assigned to the boundary limits of original cross-sections of the Halff model.

In a dam break event, there could be considerable amount of turbulence and entrainments of debris for many miles downstream of the breached section. In addition, a dam break flood, potentially with entrained debris, could overflow the river banks into the floodplains as well as inhabited areas, where the roughness could be considerably higher than those under severe flood conditions such as a SPF. To account for these conditions, the Manning's n values used by Halff in its HEC-RAS model were adjusted upward conservatively by a factor of two for 4 miles immediately downstream from the each of the failed dams, i.e., 4 miles downstream from Buchanan Dam and Mansfield Dam, respectively. For the rest of the model river reach, the Manning's n values were assumed to be 1.2 times that used in the Halff study (Base Case). A sensitivity case was performed using the same Manning's n values as in the Halff study, except for a 4-mile distance downstream from Buchanan Dam as well as from Mansfield Dam where the Manning's n values were two times the values used in the Halff study (Sensitivity Case).

Increasing the Manning's n values increases the simulated water levels because of increased roughness and therefore is a conservative approach in estimating the maximum flooding water levels at the plant site.

2.3.1.2.1.5 Predicted Water Levels at STP 1 & 2 from Upstream Dam Failure Model The HEC-RAS dam breach and unsteady flow routing model (Base Case) predicted that the peak water level at the STP site, without considering the wind wave effects, due to the domino-type failure of the upstream dams would be at El. 28.6 ft MSL or 28.4 ft NAVD 88. The discharge at the time of the peak water level would be 1.87 x 106 cfs. For the Base Case, the flood wave would take about 65 hours7.523148e-4 days 0.0181 hours 1.074735e-4 weeks 2.47325e-5 months to reach STP 1 & 2 after Mansfield Dam fails. This flood wave travel time would be about 58 hours6.712963e-4 days 0.0161 hours 9.589947e-5 weeks 2.2069e-5 months for the Sensitivity Case. The predicted dam break flood and stage hydrographs for the two cases are presented in Figures 2.3-9 and 2.3-10. The simulated maximum dam break water surface profile from Buchanan Dam to the downstream boundary for the Base Case and Sensitivity Case are depicted in Figures 2.3-11 and 2.3-12, respectively.

The maximum still water level of El. 28.6 ft MSL (28.4 ft NAVD 88) predicted for the bounding case analyzed in this reevaluation, using present-day technology and the most up-to-day watershed and hydrological data available, is lower than the still water levels documented in the UFSAR for the two most critical upstream dam failure scenarios. The UFSAR estimated a still water level of El. 32 ft MSL for the postulated failure of Buchanan and Mansfield Dams (in which the Buchanan breach hydrograph was added to that for Mansfield), and a still water level of El.

34.1 ft MSL for the postulated cascade failure of all major upstream dams (Reference 2.3-1, Subsections 2.4.4.1.1.1 and 2.4.4.1.1.4). The comparison demonstrates definite conservatisms in the UFSAR upstream dam break analysis.

Dam Breaches and Failures 2.3-8

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 FukushiniaResponse Project 2.3.1.3 Maximum Water Level at the STP I & 2 Site Determination of water level at the STP 1 & 2 site with the consideration of wind and wave effects based on the predicted maximum still water level from the upstream dam failures is described in the following.

2.3.1.3.1 Water Level at the STP I & 2 Site with Wind Setup In accordance with the guidelines in ANSI/ANS-2.8-1992, Reference 2.3-7, the maximum dam breach flood level at the plant site needs to consider the wind setup and wave runup effect from the coincidental occurrence of a 2-year design wind event. The 2-year fastest mile wind speed at the site is 50 mph based on Reference 2.3-7. The methodology given by the Coastal Engineering Manual (CEM), Reference 2.3-12, was adopted along with References 2.3-13, 2.3-14 and 2.3-15 to estimate the wave height and wave run-up at STP 1 & 2 power block and Essential Cooling Water (ECW) intake structure. The procedures outlined in CEM use the wind speed, wind duration, water depth, and over-water fetch length, and the run-up surface characteristics as input. As discussed in the UFSAR for STP 1 & 2 (Reference 2.3-1), accurate estimates of the fetch length for this flooding scenario could not be made. Based on the topographic variations and any man-made features that would limit wind effects, however, two critical fetches were identified as shown in Figure 2.3-13; one in an easterly direction towards a low lying ridge, Fetch B, and the other along the Colorado River in a northeasterly direction, Fetch A. The fetch in the easterly direction, Fetch B, was estimated to be about 15.5 miles with a maximum water depth varying from 1 to 23 ft at the peak of the dam break flood. The fetch along the northeasterly direction, Fetch A, was estimated to be about 17.6 miles, with a maximum water depth varying from 1 to 9 ft at the flood peak.

Using the method based on the standard wind-stress and hydrostatic pressure balance as presented in Reference 2.3-14, the maximum wind setup at the STP 1 & 2 site of about 2.9 ft is estimated for a wind speed of 39.3 mph for Fetch A with an average depth of about 7.1 ft. The maximum wind setup induced by a wind speed of 39.5 mph over Fetch B with the average depth varying from 3.8 ft to 9.9 ft along the fetch, divided into three segments, is estimated to be 3.0 ft. It should be noted that, UFSAR for STP 1 & 2 estimated the wind setup of 1.6 ft at the plant (for the postulated failure of Buchanan and Mansfield Dams) (Reference 2.3-1, Subsection 2.4.4.3.1).

Adding to the maximum water level of El. 28.6 ft MSL, estimated by the HEC-RAS dam break model for the STP site, the maximum still water level from the dam failure flooding scenario, including the wind setup would; therefore, be at El. 31.6 ft MSL.

2.3.1.3.2 Wave Runup on Power Block Buildings With the surrounding site grade around the power block at El. 28.0 ft MSL (Reference 2.3-1, Subsection 2.4.1.1), the maximum water depth at the STP 1 & 2 power block, including the wind setup, is about 3.6 ft, during the upstream dam break event. For this shallow water depth, a breaking wave condition prevails. Assuming that the incident wave direction is perpendicular to the buildings, a conservative assumption, the maximum wave height, peak wave period at the STP site and the resulting wave runup on the safety-related buildings are estimated to be 3.1 ft, 4.5 sec and 4.7 ft, respectively, using on Goda's formula given in Reference 2.3-13, that utilizes the maximum wave height and direction to determine the maximum wave runup on a vertical Darn Breaches and Failures 2.3-9

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project wall. As a conservative assumption, the incident wave is assumed to approach the ECW intake structure, perpendicularly. Therefore, the maximum water level with wave runup is expected to reach at El. 36.2 ft MSL on the buildings, including the safety-related structures.

The water level prediction, including wave runup, from the reevaluation of the bounding upstream dam failure case as illustrated above, is over 7 ft lower than the corresponding maximum flood levels resulting from the two most critical upstream dam failure events as documented in the UFSAR. The UFSAR for STP 1 & 2 (Reference 2.3-1, Subsection 2.4.4.3.1) predicted a wave height and runup of 4.3 ft and 9.8 ft, respectively, resulting in a maximum water level of El. 43.4 ft MSL at the plant structures for the postulated failure of Buchanan and Mansfield Dams (in which the Buchanan breach hydrograph was added to that for Mansfield).

For the postulated cascade failure of upstream dams, UFSAR (Reference 2.3-1, Subsection 2.4.4.3.3) reported the predicted wave height of 4.8 ft and the maximum water level at the plant structure of El. 43.7 ft MSL. The lower water level obtained using present-day methods and data in this reevaluation demonstrates that the upstream dam break analysis conducted for the UFSAR was very conservative with a considerable margin embedded in the results.

The difference in the maximum water levels estimated in the UFSAR for STP 1 & 2 and the current reevaluation is mainly due to difference in the still water level (without wind-wave actions), which is predicted to be El. 28.6 ft MSL in the current reevaluation, lower than the corresponding levels of El. 32 ft MSL and El. 34.1 ft MSL in the UFSAR for STP 1 & 2 as explained in Subsection 2.3.1.2.1.5.

2.3.1.3.3 Water Level at ECW Intake Structure Figure 2.3-14a shows the layout of Essential Cooling Pond (ECP) and the ECW intake structure. The maximum wind setup at the toe of the ECP embankment outer slope is estimated to be 3 ft, based on a grade elevation at the toe, crest width and outer slope of the ECP embankment of El. 27.0 ft MSL, 6 ft and 3(H):1(V), respectively (Figure 2.3-14b).

The significant and maximum wave heights at the toe of the ECP are estimated to be 2.8 ft and 3.7 ft, respectively. Such waves would result in a maximum wave runup of 11.8 ft, exceeding the ECP embankment crest elevation of El. 34 ft MSL. Thus, the waves would overtop into the ECP.

Figure 2.3-15 depicts a schematic of wave overtopping process on an embankment.

The wave overtopping analysis is performed using the formula given in Reference 2.3-15 and conservatively assuming irregular incident waves, perpendicular to a smooth and impermeable slope. The formula utilizes spectral wave height and period, and the embankment slope and freeboard to estimate wave overtopping rate on an emerged embankment. An average wave overtopping rate of about 1,300 cfs over the ECP embankment is estimated during the peak stage of the dam break flooding at the site, assuming that the incident wave is approaching perpendicular to approximately 2,500 ft long segment of the ECP embankment that is exposed to the direction along the more critical fetch (Fetch A) (Figure 2.3-14a). Further, a conservative estimation of the ECP volume of approximately 1,600,000 ft3 above the minimum operating water level of El. 25.5 ft MSL is used, which assumes that the water surface area at the crest level (El. 34 ft MSL) is 46.5 acres, same as that at El. 25.5 ft MSL (Reference 2.3-1, Subsection 9.2.5). Under these highly conservative assumptions, the analysis shows that such a wave overtopping rate would fill up the ECP in about 20 minutes. Therefore, the initial condition for the following wave runup evaluation assumes conservatively that the ECP would be at full pool (up to the crest of the embankment) which will produce the highest runup level.

Dam Breaches and Failures 2.3-10

Enclosure NOC-AE-13002975 FloodingHazardReevaluation Report STP I & 2 Fukushima Response Project To determine the wave runup on the ECW intake along Fetch A, which is more critical than Fetch B, the transmitted wave height into the ECP is calculated, assuming that the water level in the ECP is equal to the ECP crest elevation (El. 34 ft MSL). The transmission coefficient of 0.33 is estimated based on the embankment crest width (6 ft), wave runup of 11.8 (ft) and freeboard height of (2.5 ft). The maximum wave height and peak wave period of the transmitted wave are estimated to be 1.2 ft and 4.5 sec, resulting in a wave runup of about 1.8 ft on the ECW intake structure. This wave runup results in a maximum water elevation of approximately El. 35.8 ft MSL at the intake structure for the bounding upstream dam failure case. The dividing dike in the ECP (Figure 2.3-14a) has a crest elevation of El. 38 ft MSL (Reference 2.3-1, Subsection 2.4.8.2.3). The dike would partially block the wave energy, reducing the height of waves approaching the ECW intake structure. However, such effects are disregarded, conservatively.

It should be noted that, the UFSAR for STP 1 & 2 estimated the transmitted wave height of 3.2 ft into the ECP and higher runup levels at the ECW intake structure for the two most critical upstream dam failure scenarios. According to the UFSAR, wave runup could reach a level of El.

39.3 ft MSL at the ECW intake structure for the case of postulated failure of Buchanan and Mansfield Dams (Reference 2.3-1, Subsection 2.4.4.3.1), and El. 38.3 ft MSL for the postulated cascade failure of upstream dams (Reference 2.3-1, 2.4.4.3.3). Comparing to the reevaluation results, the UFSAR flood levels predicted at the ECW intake structure for the two most critical upstream dam failure events are higher by at least 2.5 ft and are therefore more conservative.

Table 2.3-5 and Table 2.3-6 summarize the results of the upstream dam break water level analyses for the safety-related structures at the STP 1 & 2.

2.3.1.3.4 Associated Flooding Impacts from Upstream Dam Failures As described above, both the still water level and maximum water levels (including wind setup and wave runup) predicted in this reevaluation for the bounding upstream dam failure case are lower than the corresponding levels documented and evaluated in the UFSAR. Therefore, there will be no additional impact on the safety functions of the plant due to inundation from the upstream dam failure flood waves.

The hydrostatic and hydrodynamic forces on the structures during an upstream dam failure event, though not specifically analyzed in the UFSAR, are bounded by the design basis flood event for the plant, i.e., the breaching of the MCR embankment, which would produce flood elevations that vary between El. 44.5 and El. 50.8 ft MSL at the power block structures and El.

40.8 ft MSL at the ECW intake structure, as described in Subsection 2.3.2. The lower flood levels predicted in the reevaluation analysis on the upstream dam failure flooding mechanism further supports this conclusion.

The potential for waterborne missiles reaching the STP site due to upstream dam failures is not considered to be a risk on account of the shallow flood depth at maximum still water of El. 28.6 ft MSL (approximately 6 inches above the nominal grade at the site of El. 28 ft MSL). In addition, the site is located in the floodplain of the Lower Colorado River where the flood flow velocities are in general substantially lower than that in the main channel and the potential for waterborne missiles does not exist. The shallow flow depths and low flow velocities on the floodplain would not produce any significant erosion and sedimentation that would impact the safety of the plant.

Similarly, debris is not considered a risk for potential adverse impact to plant safety due to the very shallow inundation flood depth at the power block. In the ECP, the operation of the ECW Dam Breaches and Failures 2.3-11

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STP1 & 2 Fukushimna Response Project intake structure will be protected by the embankment crest (at El. 34 ft MSL) which is over 5 ft higher than the predicted still water level of El. 28.6 ft MSL. In the unlikely event that some debris may be carried into the ECP with the wave overtopping flow during the bounding upstream dam break event, no adverse impact to the ECW system is expected as the intake is protected by trash bars and traveling water screens which were designed to filter out debris to protect the operation of the ECW pumps. Further, any impact from debris and sediment on the plant, especially on the ECP and ECW intake structure, would be bounded by the MCR embankment breach design basis flood mechanism due to the proximity of the MCR to the plant facilities.

2.3.2 MCR Breach Evaluation The reevaluation of the flooding hazards from the onsite Main Cooling Reservoir (MCR) embankment failure event is described in this subsection. As illustrated in the UFSAR of STP Units 1 & 2, this flooding mechanism constitutes the current design basis for both the safety related structures and facilities, and those that are important to safety, in the power block, the Essential Cooling Pond (ECP) and the Essential Cooling Water Intake Structure (ECWIS).

The flooding impact as a result of a MCR embankment breach event was analyzed in details as documented in the UFSAR for STP 1 & 2 (Reference 2.3-1, Subsection 2.4.4). The UFSAR MCR embankment breach analysis use three hydrodynamic models, as summarized in following subsections, to simulate the flood waves from different breach configurations and the resulting flood levels within the model domain. This reevaluation effort considered the hierarchical hazard assessment (HHA) and includes a comprehensive review of the UFSAR MCR breach model study to confirm that the approach and methodology, modeling tools, supporting data and results meet the present day requirements specified in NRC 50.54(f)

Request for Information (RFI) letter of March 12, 2012. In addition to the numerical modeling software used, input data such as model grid, initial and boundary conditions, and formulation of the breach, conformance with regulatory guidance and industry standards are reviewed.

This reevaluation also includes results from the MCR embankment breach evaluations performed in support of the STP 3 & 4 Combined License Application (COLA) (Reference 2.3-2) to examine qualitatively the margin available in the UFSAR current design basis flood level and the sediment impact on the power block. An additional flood analysis study conducted for STP 1

& 2 (Reference 2.3-19) is incorporated in this reevaluation to describe the debris and sediment impacts associated with the MCR breach event on the ECP and ECWIS.

Review of the UFSAR MCR breach evaluation, supplemented with other more recent MCR related flood analyses, indicates that the current flood design basis as presented in UFSAR for STP 1 & 2 is valid and remains bounding, and the two-dimensional MCR embankment breach model study conducted in support of the development of the design basis flood elevation as documented in Subsection 2.4.4.2.2 of the UFSAR meets the flooding reevaluation objectives of using present-day methodology and data. In particular, the analysis follows the guidance of ANSI/ANS 2.8-1992 (Reference 2.3-7) on the assessment of the potential dam failure modes and the specification of the antecedent and combined event conditions, consistent with the recommendations of NUREG/CR-7046. In addition to demonstrating that the current flood design basis, which is controlled by the MCR breach flooding event, provides the bounding flood levels for the plant, the assessment also indicates that there are potential margins available.

Daia Breaches and Failures 2.3-12

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project 2.3.2.1 Summary and Review of the UFSAR for STP I & 2 MCR Breach Analyses A summary of the MCR breach analyses presented in the UFSAR for STP 1 & 2 (Reference 2.3-1) is presented below together with a review of the analyses.

2.3.2.1.1 Summary of the UFSAR for STP 1 & 2 MCR Breach Analyses The MCR is enclosed by a rolled-earthen embankment, rising an average of 40 ft above the natural ground surface south of the plant site with surface area of approximately 7,000 acres. The centerline of the north embankment is approximately 800 ft south of the centerline of the power block of STP 1 & 2 (Reference 2.3-1, Subsection 2.4.4). Site grade near the northern embankment is in the range of El. 27 ft MSL to El. 29 ft MSL, and the top of the embankment is at about El. 65.75 ft MSL (Reference 2.3-1). Normal maximum operating level of the reservoir is at El. 49.0 ft MSL (with approximately 202,700 acre-ft capacity), which is about 20 to 22 ft higher than the site grade near the northern embankment (Reference 2.3-1, Subsection 2.4.4).

Postulated failure mechanisms of the earthen embankment includes excessive seepage from piping through the foundations of the embankment, seismic activity leading to potential liquefaction of the foundation soils, and erosion of the embankment due to overtopping from flood or wind-wave events. However, as indicated in the UFSAR for STP 1 & 2 (Reference 2.3-

1) failure of the MCR embankment due to any of these probable mechanisms is not considered a credible event, because of design adequacy of the MCR embankment and regular monitoring and maintenance program that would prevent potential embankment failure. Nevertheless, a conservative approach was adopted in the flood risk evaluation to assume that the embankment would fail. The most conservative conjecture of such a failure suggested that an embankment section of several hundred feet long would translate downstream several tens of feet off of its original location. This failure scenario was modeled using a 2-dimensional hydrodynamic model by assuming an instantaneous removal of a 400-ft long section of the embankment. In order to ensure sufficient freeboard in the design of the safety related facilities for flood protection, the postulated breach length was further increased from 400 ft to 4000 ft, incrementally, to determine the most critical flooding impact to the site. An approximately 2000 ft or wider breach was found to produce the highest flood level at the safety facilities of STP 1 & 2.

In addition, the model analysis of the postulated failure of the MCR embankment presented in the UFSAR of STP 1 & 2 included other highly conservative assumptions:

1. Although historical embankment and dam failure events typically involve a time lapse from the onset of failure to full development of breach, an instantaneous removal of the section of the embankment was conservatively adopted instead of using a more realistic rate of breach development.

2. The embankment breach analysis takes no credit for flow retardation and dispersion that, in reality, would be provided by the circulating water intake pipes, Circulating Water Intake Structure (CWIS), Circulating Water Discharge Structure (CWDS), and various other obstructions between the embankment and plant structures.

Three hydrodynamic models were developed for the MCR breach flooding assessments. Two of the models (Model No.1 and Model No. 2) were developed by using a two-dimensional finite difference computer program known as "System 21" developed by Danish Hydraulics Institute (Reference 2.3-18a) to assess flood levels at the structures within the power block area and also on the ECWIS. The third model (Model No. 3) was developed by using National Weather Dam Breaches and Failures 2.3-13

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Service (NWS) DAMBRK program (Reference 2.3-18b) to assess flood wave impact on the ECP southern embankment. The NWS DAMBRK program was also used to conduct a design confirmation of the DHI model results. Analysis details and results from the three models are summarized below.

Numerical Modeling Codes:

The DHI "System 21" computer program (Reference 2.3-18a) used for Models 1 and 2 solves the partial differential equations for conservation of mass (volume) and linear momentum under the assumptions of nearly-horizontal two-dimensional flows. The equations account for nonlinear effects in shallow water and comprise terms for topographical irregularities and for bed shear stress as given by Manning's roughness coefficient. The equations are reformulated as implicit difference equations with the variables defined on the following space-staggered rectangular grid. A five-level implicit difference scheme, utilizing a "double sweep" method, ensures high-order accuracy through the use of time and space centering of all derivatives and the most significant coefficients.

The DHI model system requires two-point initial conditions and one-point boundary conditions.

Initial conditions usually are specified as a given water surface elevation or depth of water at each grid point, and zero flow. In simulating wave-front propagation, a thin layer of water is usually assumed throughout the model outside the reservoir for purposes of computational convenience. The initial water surface at the wave front is assumed to slope over the distance of one grid spacing. This description approximates the shape of the wave front after a short period of time (assumed 1 or 2 sec).

In the wave-front region the flow is far from horizontal, and the difference scheme would not, by itself, be able to resolve the front because it is non-dissipative. To dissipate the energy that is physically lost by turbulence and short-wave radiation, a specially designed module, a dissipative interface, is employed in the computer system. This dissipative interface can be adjusted to dissipate the correct amount of energy while conserving mass and linear momentum. In Models 1 and 2 the dissipative interface was adjusted to produce a slightly steeper wave front than would occur physically and, therefore, predicts impact water levels which are conservative.

The NWS's DAMBRK computer program is used to develop the outflow hydrograph from a dam and hydraulically route the flood through the downstream valley. The governing equations of the model are the complete one-dimensional Saint-Venant equations of unsteady flow which are coupled with internal boundary equations representing the rapidly varied flow through structures such as dams and bridge/embankments which may develop a time-dependent breach. Also appropriate external boundary equations at the upstream and downstream ends of the routing reach are utilized. The system of equations is solved by a finite-difference method.

Model Configurations:

Model No. 1 was developed to simulate flood-wave impacts (i.e., initial flood-wave runup) on the south side of the power block and the ECWIS. Figure 2.3-16 shows the coverage of Model No.

1 as well as those of Model No. 2. Simulations were conducted in an unsteady-state mode with a computational time step of 2 seconds. The model grids had a resolution of 70 by 70 ft. The maximum water level results from this model setup were lower compared to the corresponding Model 2 results. However, this model setup was used to determine the critical embankment Dam Breaches and Failures 2.3-14

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project breach length that would produce the maximum water levels at the power block and the ECP. In order to determine the sensitivity of runup-breach length relationships, several breach lengths were investigated. The shortest breaches were located so as to create maximum runup on Unit 2, while the longer breaches were placed to produce identical runup on both units.

Model No. 2 used a quasi-steady-state mode to assess the maximum water levels which would be expected several minutes after the postulated breach. Model simulations were conducted in two grids: 1) a fine grid (with 70 by 70 ft grid resolution) with modeling domain focused on the power block and the ECP areas and 2) a coarse grid (with 210 by 210 ft grid resolution) with a larger modeling domain so that model boundaries are far from the areas of interest to minimize boundary effects. Figure 2.3-17 shows the layout of coarse grid relative to the facilities in the power block and the ECP areas and Figures 2.3-18 and 2.3-19 show the layout of the fine grid model and the details at Units 1 & 2. The coarse grid model was run first to generate the downstream boundary condition for the fine grid model, which was run subsequently to generate water levels. Time steps of 2 and 6 seconds were used for the fine and coarse grid model simulations, respectively. The water level results from Model No. 2 are higher compared to the corresponding results from Model No. 1, and were used to establish the design basis flood levels for the safety-related facilities in the power block and at the ECP including the ECWIS.

Model No.3 used the NWS's DAMBRK program (Reference 2.3-18b) to assess the flood-wave impact on the ECP southern embankment. The crest elevations of the ECP embankment and the interior dike are, 34 and 38 ft MSL, respectively. In this model, all of the breached flow was conservatively assumed to go over the interior dike, which being 4 ft higher than the perimeter embankment, would cause the water surface to attain its maximum elevation.

Initial Conditions:

The initial water level in the MCR for the breach was specified based on the standard project flood (SPF) water level for the MCR, which was obtained by routing the standard project storm (SPS), 25.28 in of rainfall in 48-hr, through the reservoir. The starting water level prior to the SPS was set to the operating level of the MCR, which is 49 ft MSL. Accordingly, the initial water level for the MCR breach was obtained was 50.5 ft MSL. This water level was set as the upstream boundary condition for all the three models.

For Model No. 1, an initial water depth of 8 in (Reference 2.3-1) was specified within the modeling domain downstream of the breach, instead of dry conditions, to obtain continuity in the computation and make it possible to describe the front propagation properly. It was found that the thin layer of water specified did not affect the results of the computations. For Model No. 2, the initial water depth in the modeling domain was specified as 1 ft.

Roughness Parameter:

The surface roughness in all the three models was represented by Manning's n values.

Evaluations were made for minimum, design, and maximum values of roughness coefficient, with resulting values of 0.035, 0.045, and 0.095, respectively, by using the method presented in Reference 2.3-20 and by considering type of surface material, degree of flow irregularity, variation of flow cross sections, relative effect of obstructions, vegetation and degree of meandering that are related to surface roughness. It was determined that the presence of obstructions represents the greatest variable with respect to the choice of this value. Major structures, including the CWIS, CWDS, central dividing dike, storage tanks, and miscellaneous Dam Breaches and Failures 2.3-15

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project buildings, were ignored in these in the analyses for conservatism. The net effect of such obstructions would be to impede flow, forcing it to disperse around plant structures and thus reduce the flood wave runup. A Manning's n of 0.046 was used in all the three models in the analyses of the MCR embankment breach flooding.

Boundary Conditions:

The upstream model boundary for the three models corresponds to the initial water level of the MCR prior to embankment breach (i.e., 50.5 ft MSL).

For Model No. 2, the downstream model boundary was placed far away from the power block and the ECP areas not to influence the maximum water levels at these areas. For Model No. 2, the downstream boundary of the coarse grid was initially held at 1 ft above ground level. Then a determination was made of the natural depth that would actually pass through the boundary if the elevation there was not kept at 1 ft. Based on the natural depth a sensitivity analysis was conducted to determine that the boundary condition had insignificant influence on the flow conditions at the power block and the ECP areas.

For Model No. 3, the upstream boundary was placed sufficiently far from the ECP such that it does not significantly influence the maximum water surface elevation during the passage of a breached flow over the interior dike. The downstream boundary was selected at the interior dike. For the downstream boundary condition, a rating curve based on discharge characteristics over an embankment shaped weir was used.

Modelinq Results:

Model No. 1 predicted an instantaneous maximum water level (wave runup) of El. 50.2 ft MSL on the south face of the power block structures. The model simulated the instantaneous and total removal of a 2030-ft section of the MCR embankment. The hydrodynamic forces resulting from this flood wave surge were found not to be the controlling factors for the flood-related loads on the structural design of the facilities.

Model No. 2 predicted maximum water level of 50.8 ft MSL at the power block area (occurring at the south face of the fuel-handling building) and 40.8 ft MSL for the ECWIS. This was for the instantaneous and total removal of a 1890-ft section of the embankment. Its water levels on the south face of the power block structures were slightly higher than the results of Model No. 1.

The maximum water elevations and buoyancy elevations (that correspond to quasi-steady-state condition) for the plant structures and the ECWIS are shown in Table 2.3-7 and Figure 2.3-20 shows the profile of the maximum water level along north-south section of the power block. As Table 2.3-7 indicates, the maximum water levels at the plant structures and at different faces of the structures are not necessarily the same because of location in regards to flood wave propagation. The flood-related hydrostatic and hydrodynamic loads for the structural design of the facilities were based on the results of Model No. 2 (Reference 2.3-1, Section 3.4).

The results from Model No. 3 were not used to establish the design basis flood levels for the power block structures or the ECP.

Dam Breaches and Failures 2.3-16

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project 2.3.2.1.2 Review of the UFSAR for STP 1 & 2 MCR Breach Analyses As described earlier, the DHI "System 21" computer program was used to simulate the MCR breach flood wave propagation and the resulting maximum water levels at the plant structures and ECWIS. The computer program which solves the 2-dimensional unsteady-state shallow water equations is considered a reasonable and appropriate numerical tool for simulating flood waves from dam failures including the flooding from the MCR embankment breach events.

A large and coarse grid model was adopted as part of Model No. 2 to include sufficient area so that approximations of downstream boundary conditions would have negligible effect on simulated flow characteristics at plant structures. The results of the coarse grid were used in a fine grid model of Model No. 2 in which flow characteristics around the plant structures and the ECP were studied in detail. For Model No. 1, a smaller domain used compared to Model No. 2, because only the instantaneous maximum water level (flood wave runup) on the plant structures was determined. As described in the UFSAR, the results from Model No. 2 produced higher maximum water levels at the plant structures.

Manning's n was selected based on Reference 2.3-20, by considering surface type and different flow conditions and characteristics. This approach of selecting Manning's n is acceptable in the present day applications of characterizing flow roughness parameters downstream of dam failures.

All secondary structures surrounding the main buildings and other structures such as CWIS, CWDS, storage tanks and other miscellaneous buildings were omitted in the models so that the most critical impact on the main structures could be obtained.

The approaches used in determining the maximum flood levels due to the MCR breach in the UFSAR for STP 1 & 2 are highly conservative for the following reasons:

" The postulated failure of the MCR embankment is not a credible event given that adequate engineering and construction considerations are given for various potential failure mechanisms during the design and construction phases, in addition to the implementations of regular monitoring and maintenance programs. Any potential failure of the embankment would be averted before it occurs.

" A extremely conservative embankment breach width of approximately 2000-ft was adopted in establishing the design basis flood levels. This is larger by an order of magnitude than the breach openings expected from past experience with dam and embankment failures, as outlined in Reference 2.3-1.

The assumed instantaneous removal of the failed portion of the embankment is a highly conservative assumption given that a complete embankment failure takes some time, in the order of tens of minutes.

In addition, debris from embankment failure, which mostly comprises of the embankment material would retard breach flow with the effect of reducing the maximum water level at the power block and the ECP areas.

The upstream model boundaries were specified to correspond to a constant initial water level of the MCR (50.5 ft MSL). The use of a static water level boundary condition is Dam Breaches and Failures 2.3-17

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushimna Response Project highly conservative, in that the water level at the breach is expected to drop with time with the reservoir storage is being depleted during an actual embankment failure event.

The effects of sediment erosion/scouring and deposition were not specifically modeled in the UFSAR but were qualitatively assessed. In addition, the potential of these effects are assessed based on the MCR embankment breach analyses conducted recently, as described in the subsequent subsections.

2.3.2.2 Comparison of the UFSAR STP I & 2 MCR Breach Analyses with Recent Analyses A recent MCR breach model study was conducted for the proposed STP 3 & 4, to be located adjacent to STP 1 &2, as part of a Combined License Application (COLA) (Reference 2.3-2) to establish the design basis flood levels and evaluate the sediment impacts at the new units. The STP 3 & 4 COLA analysis assumes a conservative but more realistic breaching scenario than the UFSAR models. Independently, another MCR breach model using the Delft3D model was developed to examine the conservatism in the flood levels at the site of the future units assuming a highly conservative breaching configuration as used in the UFSAR. Finally, a recent MCR breach analysis, which used a more realistic breaching scenario than that used in the Units 3 & 4 COLA model, was conducted in 2012 for STP 1 & 2 to acquire a better understanding of the site conditions that would be expected following the very unlikely event of a failure of the Main Cooling Reservoir embankment. In addition to considering more realistic water elevations and velocities, this analysis considered the types of debris that might be produced by an embankment failure and where the debris might be conveyed, particularly at the ECP and ECWIS (Reference 2.3-19). The results from these recent models are evaluated to assess the conservatisms in the design basis flood levels derived from the MCR breach analysis described in the UFSAR. The three model analyses are summarized below. Figure 2.3-21 (from Reference 2.3-2) shows the location of the proposed STP 3 & 4 in relation to STP 1 &

2.

2.3.2.2.1 Summary of Recent MCR Breach Analyses MCR Breach Analysis of COLA FSAR for STP 3 & 4, Subsection 2.4S.4:

The COLA FSAR for STP 3 & 4, Subsection 2.4S.4 (Reference 2.3-2) determined the MCR embankment breach width based on empirical breach parameter relationships that were described and summarized by the Dam Safety Office of the U.S. Bureau of Reclamation (Reference 2.3-21). The most critical breaching location with respect to flooding at STP 3 & 4 was considered.

The breach bottom width was estimated based on Froehlich's equation presented in Reference 2.3-21 and was selected because it provides the largest breach width estimate among the various methods evaluated. In addition, the equation predicts an average breach width of 220 m (722 ft) for the Teton Dam, although the actual average breach width of Teton Dam at failure was only 151 m (495 ft) indicating the conservatism of the equation. Based on the equation, the average breach width of the embankment is 417 ft. Given that the trapezoidal geometry of the breach, an average breach width of 417 ft corresponds to a bottom breach width of 380 ft for the assumed side slopes of 1 vertical to 1 horizontal. This ratio was obtained from field observations for earth-filled structures presented in Reference 2.3-21.

Dam Breaches and Failures 2.3-18

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushimna Response Project Time to fail of the embankment was based on the equation given by MacDonald and Langridge-Monopolis as presented in Reference 2.3-21. The equation was found to predict a time to fail that came closest to describing a breach expansion rate meeting that of Teton Dam. Breach expansion rate is determined using the predicted breach width and the time to fail. As indicated in Reference 2.3-21, the time from beginning of rapid growth of breach to stopping of significant lateral erosion process at Teton Dam was estimated at 1.25 hours2.893519e-4 days 0.00694 hours 4.133598e-5 weeks 9.5125e-6 months and the final breach width was 496 ft, resulting in an expansion rate of 198 feet per hour. This rapid rate of erosion was considered to be due to the higher hydraulic depth to drive the outflow and associated erosion.

The MacDonald and Langridge-Monopolis equation predicted a time to fail of 1.7 hours8.101852e-5 days 0.00194 hours 1.157407e-5 weeks 2.6635e-6 months for the MCR embankment. This gave a breach expansion rate of approximately 112 feet per hour, which is smaller than the Teton Dam and was considered acceptable because of lower hydraulic depth of the MCR compared to the Teton Dam.

The modeling program FLDWAV, developed by the National Weather Service (Reference 2.3-22), was used to generate the outflow flood hydrograph from the MCR embankment breach, based on breach parameters discussed earlier. FLDWAV is a parametric numerical model that can generate breach outflow hydrograph based on user-defined input of breach parameters.

FLDVAV predicted a peak flow of 130,000 cfs at 1.7 hrs after commencement of embankment failure. The flood hydrograph is shown in Figure 2.3-23 and was used as input to RMA2, to model flow downstream of the breach. RMA2 is a two-dimensional depth-averaged finite-element hydrodynamic numerical model developed by the United States Army Corps of Engineers (USACE) (Reference 2.3-23). The computer program can simulate dynamic water surface elevations and horizontal velocity components for subcritical, free-surface flow in a two-dimensional flow field. Figure 2.3-24 shows the layout of the modeling domain.

The Manning's n value for each surface or material type was assigned based on typical values published by the United States Geological Survey (USGS) and the HEC-RAS manual (References 2.3-24a and 2.3-24b). Each major building was evaluated on whether it would remain in place following the flood caused by a MCR embankment breach. Those buildings that were assumed to remain in place were considered "hard buildings." Any hard buildings higher than elevation 62 ft MSL were considered to block the flow, and therefore were shown as blank areas in the model. Those buildings assumed to fail were considered "soft buildings." Soft buildings were assumed to be destroyed with foundation slab remaining in the grid. These buildings were considered "high drag" areas with a higher roughness value to represent the effects of remaining frames and debris. Any buildings not included in the model were represented by a higher Manning's n value. Due to the resolution of the grid, the Vehicle Barrier System around the power block was not built into the model, but instead was represented by higher Manning's n value. Figure 2.3-25 shows Manning's n values assigned to various surface or material types represented in the model.

The downstream boundaries of the model were positioned far enough so that the maximum flood level at the STP 3 & 4 safety-related structures due to a MCR embankment breach would occur before the flood front reaches the boundaries. A constant water surface elevation was defined for the downstream boundary condition. A sensitivity analysis was performed on the downstream boundary conditions, and was indicated that the maximum flood level at the power block area is not affected by the downstream boundary conditions.

The initial water level in the MCR considered for the breach analysis was 50.9 ft MSL. This level corresponds to the response of the MCR to one-half PMP with a starting water level Dain Breaches and Failures 2.3-19

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project corresponding to the normal maximum operating level and with the addition of wind set-up produced by the 2-year wind speed (50 mph) from the south (Reference 2.3-7).

For the analysis, two critical breaching scenarios were considered, one for Unit 3 and another for Unit 4. The centerline of the reactor buildings of STP 3 & 4 is approximately 2340 ft north of the centerline of the northern embankment. The maximum water level obtained from the MCR breach is 38.8 ft MSL and the maximum velocity of the flood flow was found to be 4.72 ft per second and occurred between Units 3 and 4. Sediment erosion and scouring, and the subsequent sedimentation were also considered. Based on a bounding analysis, the sediment depths near the power block were shown to range between 0.35 and 0.40 ft. Assuming that the flood water elevation were raised by the same amount, the maximum water level due to the MCR breach flood was modified to 39.2 ft MSL.

MCR Breach Analysis usingq Delft3D Model:

The independent MCR breach analysis, that used Delft3D model, adopted the approach used in the UFSAR STP 1 & 2, in that the critical breach width that produced the maximum water level at the power block area was determined by a sensitivity analysis in which the breach width was varied between 197 ft (60 m) and 5,545 ft (1,690 m). This analysis was not used to establish the design basis flood level for STP 3 & 4. It is included here as another point of reference to demonstrate the conservatism in the current flood design basis for STP 1 & 2. The model results indicated that an instantaneous breach opening, with a width of 4,757 ft (1,450 m) and centered along STP 3 & 4 centerline would produce the maximum water level of 47.9 ft MSL at the power block (at the south face of STP 3 & 4 Ultimate Heat Sink). Similar to the UFSAR for STP 1 & 2 analysis, a Manning's n value of 0.046 was used throughout the modeling domain and small structures, including building, between the MCR and the power block area were assumed to be washed away during an instantaneous MCR breaching event, reducing the flood impeding effect of the structures. The initial water level in the MCR prior to MCR breach was specified at 50.74 ft MSL, corresponding to one-half PMP event. Wind effects at the structures were neglected on the basis that flooding from the MCR breach is not sustained long enough for wind effects to be significant. The analysis was conducted with the Delft3D-FLOW hydrodynamic modeling program (Reference 2.3-18), in a two-dimensional domain, and included the whole MCR with the internal dikes removed for conservatism. The model boundaries were place far away from the power block area so that the peak water level is not influenced by the boundary conditions. Figure 2.3-22 shows the layout of the modeling domain.

Delft3D-FLOW is a multi-dimensional (2D or 3D) hydrodynamic (and transport) program which simulates unsteady flow and transport phenomena that result from tidal and meteorological forcing on a rectilinear or a curvilinear boundary fitted grid length. Delft3D-FLOW solves the Navier-Stokes equations for an incompressible fluid, under the shallow water and the Boussinesq assumptions. The set of partial differential equations, in combination with specified set of initial and boundary conditions is solved on a finite difference grid of model domain.

For rapidly varying and supercritical flows, Delft3D-FLOW uses a "Flooding scheme" that can be applied to dam/embankment breach and torrential river flows in vertically average 2D form. The advection scheme also computes flood wave propagation speed for both dry and wet bed conditions.

Dam Breaches and Failures 2.3-20

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 FukushiinaResponse Project 2012 MCR Breach Floodinq Analysis for STP 1 & 2:

A MCR breach analysis conducted in 2012 for STP 1 & 2 used a more realistic breaching scenario than that used in the STP 3 & 4 COLA model. The objective was to acquire a better understanding of the site conditions that would be expected following the very unlikely event of a failure of the Main Cooling Reservoir embankment (Reference 2.3-19). In addition to considering more realistic water elevations and velocities, the analysis considered the types of debris that might be produced by an embankment failure and where the debris might be conveyed, particularly at the ECP and ECWIS. The study is similar to the COLA FSAR for the STP 3 & 4 (Reference 2.3-2), as applied to STP 1 & 2, with a similar modeling approach and the use of the hydrodynamic model RMA2. The major differences are that the breaching hydrograph was generated with the BREACH model instead of FLDWAV, model grid details and location of embankment breaches.

BREACH is a physically based mathematical model developed by the National Weather Service used to predict the breach characteristics (size and time of formation) and the corresponding outflow hydrograph (Reference 2.3-25). The main parameters used are dam embankment material properties; particle size D50, unit weight, friction angle, and cohesive strength. The model predicated a peak flow of 83,200 cfs at 6.5 hrs with breach bottom width of 361 ft, and the final breach width reached 448 ft after 30 hrs. Figure 2.3-23 shows the flood hydrograph from BREACH model.

The same breach hydrograph from BREACH was also considered as part of a confirmatory analysis to verify the conservatism of the FLDWAV in the COLA FSAR for STP 3 & 4 (Reference 2.3-2), which also includes details of the on the BREACH model.

The initial water level at the MCR for the BREACH model, 50.9 ft MSL, was based on the one-half PMP and 2-yr fastest-mile wind speed from the critical direction (south) that would produce the maximum wind setup. The downstream boundary condition was specified the same way as the COLA FSAR for STP 3 & 4 analysis described earlier. The Manning's n values specified for the different surface types are shown in Figure 2.3-26 and the approach used is the same as in COLA FSAR for STP 3 & 4, as described earlier.

Three different breach scenarios based on breach locations were considered for the study with the first two located directly south of Unit 1 and 2 facilities and the third one located at the southwest of the Nuclear Support Center (NSC) Building. The locations were selected to evaluate the MCR breach impact on the two units and the ECP, respectively. The three model configurations, indicating the locations of breach, are shown in Figure 2.3-27a through Figure 2.3-27c. The flood hydrograph monitoring points used in the analyses are indicated in Figure 2.3-28.

Summary of the maximum velocities, water depths and water surface elevations is provided in Table 2.3-8 for the three simulation cases: Unit 1, Unit 2 and NSC. The maximum water elevation at the safety-related structures of the power block is 42.5 ft MSL occurring on the southern face of Unit 1 (the fuel-handling building) and the maximum water level at the ECWIS is 35.4 ft MSL. Both maximum water levels are related to the MCR embankment breach south of Unit 1.

The maximum velocity vectors for each scenario are indicated in Figure 2.3-29a to Figure 2.3-29c and the time history at each monitoring points for each scenario are indicated in Figure 2.3-Darn Breaches and Failures 2.3-21

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 30a to Figure 2.3-30c. As Table 2.3-8 indicates, higher velocities occur at Points 1 to 5, which are closest to the MCR embankment, as would be expected. The locations with high velocities have a gravel surface as indicated in Figure 2.3-25. In addition, the high velocities are only sustained for a limited period, i.e., for a few hours. There are differences reflecting the blocking effect of the units. For example, with the Unit 1 breach the highest velocities are at Points 1, 3, and 4, while point 2, directly opposite the breach, has a lower peak velocity. This reflects flow being blocked and diverted around Unit 1. A similar situation exists with the Unit 2 breach where the peak velocity at Point 1 (which is directly in the path of the breach but blocked by Unit 2) is lower than at Point 2 (which is off to the side). In the immediate area of the MCR embankment, the velocity values are greater than 10 ft/s which could damage lighter structures (as assumed) and also cause significant scour. However, scour hole from the MCR breach is not expected to reach the plant structures, as discussed in Subsection 2.3.2.4.

The blocking effect can also be seen in the water depth results. The highest water depths are at the points immediately in the breach path where the flow is blocked by the reactor units. Point 2 is in the path of the Unit 1 breach and has a depth of 14.2 ft, and Point 1 is in the path of the Unit 2 breach with a peak depth of 14.0 ft. Peak depths are lower away from these points.

2.3.2.2.2 MCR Breach Analyses Evaluation Summary A comparison of the UFSAR for STP 1 & 2 MCR breach analysis with the three recent MCR breach model analyses is provided below. The comparison shows that the current design basis flood levels (at the power block and ECW intake structure) are conservative and have potential margins available.

As discussed in Subsection 2.3.2.1, the MCR instantaneous breach width of about 2000-ft adopted in the UFSAR for STP 1 & 2 is highly conservative. Similarly, the independent Delft3D analysis for the MCR breach, also adopted a conservative breach width and time to failure. The MCR breach analysis for STP 3 & 4 COLA, showed that a smaller breach geometry, with bottom width of 380 ft and side slope of 1V:IH, would be a realistic estimate. This was supported based on historical data on dam breaches, such as Teton Dam and also on formulations of dam breach parameters that are currently in use. In addition, a time to failure of 1.7 hours8.101852e-5 days 0.00194 hours 1.157407e-5 weeks 2.6635e-6 months would be a more realistic estimate than the instantaneous breach time assumption. This was also supported with the same historical data and corresponding dam breach formulations. The results from the STP 3 & 4 COLA are not directly applicable to STP 1 & 2, because the proposed STP 3 & 4 is farther away from the MCR embankment compared to STP 1 & 2, by approximately more than 1,000 ft. Therefore, a qualitative comparison was made between STP 3 &4 COLA and the independent MCR breach analysis results, so that the effect of adopting a smaller and realistic breach width is assessed.

The maximum water level at the power block area from the independent MCR breach analysis with the instantaneous MCR embankment breach width of 4,757 ft is 47.9 ft MSL. The corresponding maximum water level from the STP 3 & 4 COLA, with average breach width of 417 ft and breach time of 1.7 hours8.101852e-5 days 0.00194 hours 1.157407e-5 weeks 2.6635e-6 months is 38.8 ft MSL. The difference between the two approaches amounts to 9.1 ft. This confirms that adopting a realistic MCR breach width and time to failure would produce a much lower maximum flood levels at points of interest. Therefore, it is expected that the maximum flood levels for STP 1 & 2 would be lower than the current design basis flood levels if a realistic MCR breach width and time to failure were adopted. This can further be demonstrated based on the recent 2012 MCR breach analysis conducted for STP 1 &

2 (Reference 2.3-19). The analysis indicates that the maximum water levels at the power block Dam Breaches and Failures 2.3-22

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPJ & 2 Fukushima Response Project structures and the EWC intake structure are 42.5 ft MSL and 35.4 ft MSL, respectively. These values are considerably lower than the corresponding maximum water levels presented in UFSAR for STP 1 & 2, 50.8 ft MSL and 40.8 ft MSL, respectively. The difference in maximum water level between the two analyses is 8.3 ft for the power block structures and 5.4 ft for the ECW intake structure. Therefore, it can be concluded that the design basis flood levels presented in the UFSAR for STP 1 & 2 are based on very conservative analyses and there are significant margins between the selected design basis flood levels and the maximum water levels expected during an actual MCR breach event.

2.3.2.3 Debris and Waterborne Missiles The UFSAR for STP 1 & 2 includes an assessment of water-borne missiles due to the MCR embankment breach (Reference 2.3-1, Section 3.5). Based on a maximum velocity of 20 f/sec, corresponding to the conservatively calculated maximum water velocity resulting from a postulated failure of the MCR embankment (Reference 2.3-1, Section 2.4), it was concluded that the effects from missiles and waterborne debris such as automobiles, utility poles, wooden planks, etc. are considerably less severe than the effects of the postulated tornado missiles.

The recent 2012 MCR breach analysis for STP 1 & 2 (Reference 2.3-19), does present a detailed assessment of the impact of debris (which also includes waterborne missiles). The details of the analysis and assessment are provided below.

The assessment of debris impact included detailed study on the potential movement and impact of various types of debris caused by the MCR breach. The debris items considered are listed below:

" Single cab pick-up

" 4 door sedan

" White rock found in the power block area

" Sheet metal from buildings

" Paper

" Chairs

" 5 by 5-ft tables

" Cubicle walls

" Miscellaneous garbage

" Mobile office trailers The debris transport was modeled by considering four flow paths that originate from the breach locations and movement of debris based on mechanics of sediment transport. The following factors in the movement of debris were considered:

" Volume/weight (related to gravitational force and submerged surface area)

" Buoyancy force Dam Breaches anid Failures 2.3-23

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project

Period of buoyancy (before loss of buoyance due to damage by objects/debris)

Stability of debris in holding together

- Drag resistance from water

- Bottom frictional resistance for negatively buoyant debris The flow paths were determined based on the output from the RMA2 simulation for the three MCR breach scenarios. The outputs include water depth, elevation and velocity at ten points along each flow path considered. Output was obtained at 3-minute interval for the entire 35-hour simulation.

The debris simulation was performed to determine where each debris item can be expected to be carried during the three MCR breach scenarios. Each simulation is a linear process taking place along a flow trace provided by the RMA2 model. A spreadsheet was set up for each flow path so that the location of an individual item can be calculated along each path as a simulation proceeds. In order to better track the movement of the debris items, the debris simulation was performed in time steps 0.3-minute. The RMA2 results were interpolated to generate data at this smaller time interval.

Table 2.3-9 presents the different types of debris material grouped by major characteristics and source. Details of the grouped debris material are given below.

Embankment Materials: The first items considered are the embankment materials consisting of clay and soil cement pieces. This material will be generated early in the breach scenario and will be essentially gone four hours after the breach flow passes its peak. Three types of material are considered. The embankment is constructed of clay, and clay particles (e.g., particles < 2 micron) tend to have very low settling rates. Because they will tend to be carried with the water outside the study area, they will not be simulated. However, as the embankment washes out some of the clay material will break off as chucks. As this material is carried in the flow, hydraulic dredging experience indicates that the material will become rounded into clay balls with a range of sizes. For the analysis it was assumed that a size range of clay balls is from 0.1 to 1.0 ft, and that the specific gravity of the compacted clay material is 1.75. The other component of the embankment material is pieces of the soil cement liner from the interior surface of the embankment. These were assumed to range in size from 0.5 to 3.0 ft and to have a specific gravity of 2.37, but are much fewer in number than the clay balls. Both the clay balls and soil cement pieces are negatively buoyant but can be pushed by the water flow. Besides the difference in density, the soil cement pieces would not tend to be rounded and are thus assigned higher drag coefficient.

The embankment material starts being added at the start of the simulation, and stops being added, when the breach is essentially finished growing (within 11 hours1.273148e-4 days 0.00306 hours 1.818783e-5 weeks 4.1855e-6 months ). The pieces were added to the initial point in each flow path located over the embankment at a uniform rate. The size of each piece is determined by a random number (RDN) multiplying the range of the diameter. For example, with the clay balls having a size range of 0.9 ft (0.1 to 1.0), the diameter of each would be equal to 0.1 +RDN x 0.9. The RDN can range from zero to 1.0 so the size can cover the full range.

The spreadsheet model takes each particle along the path for the simulation period. The end product for the two embankment particle items is a set of distances for 100 cement and 100 clay balls along a particular flow path.

Dam Breaches and Failures 2.3-24

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Vehicles: Three types of vehicles were considered. Each starts off as being buoyant and will be carried by the flowing water provided that the depth is sufficient. At each time step, the spreadsheet determines if the object is positively or negatively buoyant. The calculation of buoyancy involves both its dry weight in pounds, the bottom area subject to water buoyancy (assumed constant during the simulations) and the submerged depth at the location and time.

The water depth times the area yields a volume and a vertical force of 62.4 pounds/cubic foot. If buoyancy is positive, the object is assigned the velocity of the water less an allowance for contact objects or other debris that will tend to slow the movement to some extent. This speed reduction relative to the water is assumed to range from 0 to 30% and is set for each vehicle at the start of its simulation using a random number generator. During the simulation for that vehicle, the velocity is assumed to be at the same percentage (0 to 30%) less than the water velocity.

If the object was calculated to be negatively buoyant, then it would be subject to drag force from the water and friction from the bottom. Because it would not be moving as fast as the water, the relative motion of the water would exert the drag force. The bottom friction was calculated as the negative buoyancy times a friction factor or coefficient. This bottom friction force must be balanced with the drag force of the water to yield a velocity that would be somewhat lower than the water. In effect, the relative velocity was calculated by balancing the horizontal forces (frictional and drag forces).

At each time step, the flow and elevation change at each point. As a debris object is carried along the flow trace, it encounters new bottom elevations and velocities. It is also subject to the stability aspect. In the case of a buoyant object such as a sedan, a reduction in the area subject to buoyancy would occur with time and this is likely to involve random variations. The net effect is that some vehicles would float longer than others. Even after they become negatively buoyant, they can be carried along in the flood flow for some distance. For each vehicle type, the starting location would be selected based on the flow path and the location of the parking lots or construction activity. The actual starting distance for each simulation was selected at random within the specified range. The end point is recorded for each vehicle at the end of the simulation.

Office contents: The office contents include flat material such as dry wall and cubicle walls that can be expected to be carried some distance but to lie flat at some point. Once that happens, they are unlikely to move further. The other items, chairs and tables, would not lie flat on the bottom but would have a large projected area to the flow with relatively little negative buoyancy.

As such, they can be expected to be carried some distance.

The dry wall and cubicle wall materials were assumed to start in the vertical position from a location close to the NSC, and be carried in turbulent flow for two times steps. After that time a random binary test (coin toss) was made to determine if they lie flat and stay on the bottom, or continue to move with the current. The chairs and tables would move from water drag acting on the area (length times buoyant height) and bottom friction, much like the vehicles. The main difference is that because they are largely open, the drag coefficient is reduced.

Miscellaneous Debris: This grouping includes substances that would be carried with the water (paper, cloth and trash), sheet metal from buildings, and the white (marble) rock used as ground cover around STP 1 & 2. The material carried with the water and the sheet metal will start from a location closest to the NSC. The white rock starts from the point closest the STP 1 & 2 Dam Breaches and Failures 2.3-25

Enclosure NOC-AE-13002975 Flooding Hazard Reevaluation Report STPJ & 2 Fukushima Response Project grounds. The sheet metal is handled in the same way as the dry wall and cubical walls. The rock is subject to scour and rolling along the bottom, with a reduced bottom friction coefficient to reflect this process.

The debris analysis indicated that some of the debris items considered would be carried to near the end of their flow paths. These items would not be expected to have a major impact on the facility, but could pose impediments to the re-entry process, particularly if the water is still covering much of the area after more than 30 hours3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months . One debris type that could be deposited near the units is the soil-cement block that would be created from the failure of the embankment. These are indicated to come to rest in the immediate areas of the units, which would have to be removed eventually.

All of the simulations involving a flow path through the ECP indicated a portion of the debris would enter the ECP and be retained. The flows would initially not overtop the embankment then move over the embankment and drop below the 34 ft MSL elevation of the embankment crest. The peak elevations above the embankment are 1.4 ft for the Unit 1 breach, 1.0 ft for the Unit 2 breach, and 2.9 ft for the NSC breach. Table 2.3-10 shows the fraction of each debris type that is predicted to be retained in the ECP for the paths that cross over the ECP, based on the 100 simulations for each debris type. A substantial portion of the clay balls, pieces of rock, chairs, and tables are simulated to be transported into the ECP and retained. The results also indicate that no vehicles would be carried over the top of the south embankment of the ECP.

This is because the calculated friction force of the vehicles was sufficient to stop movement, even though the peak height of flow over the embankment (2.9 ft for the NSC breach) might be considered to be sufficient to move at least smaller vehicles over the embankment. However, even if a vehicle were carried into the ECP, the flow in the ECP would be towards the north and east, away from the intake structure.

Debris such as fine clay particles from the embankment, paper, rags and garbage would move with the water and be subjected to significant settling process. The simulations indicate that most of these items can be expected to be conveyed out of the immediate study area (past FM 521). While most of these would be out of the power block area, some would be carried into the ECP. However, no impact on the operation of the ECW system is expected because the safety-related ECW pumps housed in the ECWIS structure are protected from influx of debris by track bars and traveling water screens.

2.3.2.4 Erosion and Sedimentation The UFSAR for STP 1 & 2 does not include a specific model analyses, but provides a qualitative assessment on the impact of scour/erosion and sedimentation on the safety-relates structures due to the MCR breach. The UFSAR concluded that an adequate margin of safety can be maintained for all credible failure mechanisms (Reference 2.3-1, Section 3.4 and Subsection 2.5.6), and therefore, the mechanistic effects such as scour and erosion associated with a postulated failure of the MCR embankment need not be evaluated (Reference 2.3-1, Section 3.4). However, the STP 3 & 4 COLA does present an assessment of scour/erosion and sedimentation impacts and is therefore used qualitatively to estimate the impact on STP 1 & 2.

The details of the analysis and assessment are provided below.

The embankment material that would erode is comprised mostly of clay, with a small percentage of sand from the internal drainage system and soil cement from the interior embankment slope lining. The erosion process will also produce a scour hole downstream of Dam Breaches and Failures 2.3-26

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 Fukushimna Response Project the breach that extends below the breach bottom elevation. As indicated in Reference 2.3-2, the dimensions of this scour hole, based on laboratory results were estimated to be 20 feet deep, 203 feet long and 380 feet wide. The scour hole would contribute 1,543,000 cubic feet of clay to the flood flow. The material that would erode from the MCR embankment would contribute an additional 1,697,314 cubic feet of clay; 75,644 cubic feet of sand; and 117,562 cubic feet of soil cement. The total volume of sediment that would erode due to MCR breach would be 3,433,517 cubic feet. No additional erosion from the power block area was considered because of the relatively low flood velocities and mostly concrete or asphalt pavement or compacted stone surfacing.

The MCR embankment breach scour hole at downstream of the breach adopted in the STP 3 &

4 COLA indicates that the length of the scour hole, along the flow direction, would be approximately 200 ft. However, the plant structures of STP 1 & 2 are approximately more than 500 ft from the centerline of the northern MCR embankment (Figure 2.3-20). Therefore, the scour hole from the MCR embankment breach would not impact the foundations of the plant structures and the ECWIS.

As discussed in Reference 2.3-2, sedimentation is not anticipated to have a significant effect on the site and the maximum water level resulting from the MCR breach flood. The majority of the clay and sand loads would be suspended in the flood flow and wash away farther downstream, north of FM 521 and beyond the STP site. The soil cement lining on the interior wall of the embankment would likely enter the water as chunks or blocks as the embankment collapses, and these large concrete blocks would be carried only a short distance from the breach before settling to the bottom. The sediment loading would cease when the breach opening expansion ends; however, low-sediment flows would continue for a number of hours afterwards until the water in MCR is emptied. This continued flow period would prevent any remaining clay or sand particles from settling and would wash away any small depositions in the study area.

The assessment from the STP 3 & 4 COLA discussed above are also applicable to STP 1 & 2 if a smaller and realistic breach width and time to failure were adopted. Any sediment from the embankment or from a scour hole at the breaching location would be washed way beyond the power block area of STP 1 & 2. Because STP 1 & 2 is closer to the MCR embankment compared to the proposed STP 3 & 4, the impact of sedimentation on STP 1 & 2 would not be significant. Even if the 0.4 ft deposition of sediment estimated in the STP 3 & 4 COLA (Subsection 2.3.2.2.1) for the power block area is adopted for STP 1 & 2, the reduction in the potentially margin available in the design basis flood levels for plant structures and ECWIS is not significant. Based on the predicted water levels from the 2012 MCR breach analysis for STP 1 & 2 at the southern face of the Fuel-Handing Building of Unit 2, the northern face of the Unit 1 Diesel-Generator Building and the ECWIS, the potential margins available are more than 8 ft, 9 ft and 5 ft, respectively.

Nevertheless, because of the relative proximity of STP 1 & 2 to the embankment, the flow velocities at the power block area are expected to be higher compared to STP 3 & 4 case, as indicated earlier in the Subsection 2.3.2.4. However, because the power block area has an asphalt and gravel surfacing, and also because of the limited duration that the peak velocities would be sustained (i.e., for a few hours), erosion and scours that would affect the safety function of the STP 1 & 2 are not expected.

The STP 3 & 4 COLA considered the change in water density, due to suspended load from scouring and breached embankment material, in calculating the hydrodynamic loading on Darn Breaches and Failures 2.3-27

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project structures. With a sediment concentration of 23 kg/M 3 , a water density of 1023 kg/M 3 or 63.85 lb/ft 3 was used for the load calculations. The UFSAR for STP I & 2 does not consider change in water density due to sediments. However, because the change in water density due to sediments is relatively small (2.3 %) and because of the potential margin available in the design basis flood levels for the plant structures and the ECWIS, preclusion of water density change would not have significant impact.

2.3.2.5 Inundation Period The 2012 STP 1 & 2 MCR breach analysis also includes results of long-term inundation periods (Reference 2.3-19). The inundation period to 0.25 ft water depth was estimated by extrapolating the water level hydrographs developed for debris assessment. A summary of the resulting inundation periods is provided in Table 2.3-11. The results provide an estimate of non-accessible time at the 10 selected locations based on available RMA2 output. The inundation periods range from 22 to 66 hours7.638889e-4 days 0.0183 hours 1.09127e-4 weeks 2.5113e-5 months at the monitoring locations indicated in Figure 2.3-28.

2.3.3 Conclusions The reevaluation of flooding hazards associated with dam and embankment failures described above concludes that the highly conservative MCR embankment breaching scenario remains the controlling flooding mechanism for STP 1 & 2, consistent with the design basis flood evaluation in UFSAR. A detailed review as part of the flooding reevaluation effort indicates that the methodology adopted in the MCR breach flooding analysis documented in the UFSAR for STP 1 & 2 is comparable with present-day technical approach and modeling techniques. The UFSAR design basis flood levels, between 44.5 and 50.8 ft MSL at the plant structures (power block area) and 40.8 ft MSL at the ECW intake structure, remain bounding with substantial margins.

The 2012 MCR breach analysis conducted for STP 1 & 2 using a more realistic breach width indicates that debris and waterborne missiles would not adversely impact the safety-related facilities at the power block area and at the ECP. In addition, based on the maximum velocity predicted in the 2012 MCR breach analysis, sedimentation or scouring would also not have an adverse effect on the safety-related structures of STP 1 & 2. Although the impact of wind effects, sediment deposition and water density change due to sediments (for calculating hydrostatic and hydrodynamic forced on structures) were not considered in the UFSAR for STP 1 & 2, they are not expected to have significant impact on the margin available in the UFSAR flooding hazard analysis results of the MCR embankment breach.

2.3.4 References 2.3-1 STPEGS Updated Final Safety Analysis Report, Units 1 & 2, Section 2.4, Revision 15; Subsection 2.5.6, Sections 3.5 and 9.2, Revision 16; Section 3.4, Revision 13.

2.3-2 South Texas Project Units 3 & 4 Combined License Application (COLA), Final Safety Analysis Report (FSAR), Nuclear Innovation North America LLC (NINA), Subsection 2.4S.4 (Potential Dam Failures), Revision 07.

2.3-3 "Texas Administrative Code - Title 30, Part 1, Chapter 299", Office of the Secretary of State of Texas, provisions adopted to be effective May 13, 1986 (11 TexReg 1978).

Dam Breaches and Failures 2.3-28

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushitna Response Project 2.3-4 "Phase II - Dam Safety Evaluation Project, Task Order B - Reconnaissance investigation, Interim Report", Volume I, Freese and Nichols, Inc., August 1992.

2.3-5 "Celebration marks completion of 10-year LCRA dam project to improve public safety", Press Release by LCRA dated January 12, 2005; available at http:/www.lcra.org/newsstory/2005/damupgrade-project.html, accessed on August 31, 2007.

2.3-6 "Disaster Ready Austin: Building a Safe, Secure and Sustainable Community", City of Austin Hazard Mitigation Action Plan, 2003 - 2008, prepared by LCRA and H20, Inc., revised on August 7, 2003.

2.3-7 "Determining Design Basis Flooding at Power Reactor Sites", La Grange Park, Illinois, ANSI/ANS-2.8-1992, American Nuclear Society, July 1992. (Historical Technical Reference) 2.3-8 "Engineering Data on Dams and Reservoirs in Texas", Part Ill, Report 126, Texas Water Development Board, February 1971.

2.3-9 "HEC-RAS, River Analysis System, Version 3.1.3", U.S. Army Corps of Engineers, Hydrologic Engineering Center, May 2005.

2.3-10 "Flood Damage Evaluation Project", Chapter 1-6, Volume Il-C, Volume Il-B, Halff Associates, Inc., July 2002.

2.3-11 "Industries Regulations, Guidelines and Manual - Engineering Guidelines for the Evaluation of Hydropower Projects", Federal Energy Regulatory Commission, April 1991.

2.3-12 U.S. Army Corps of Engineers, Coastal Engineering Manual, Part I1: Chapter 1, Chapter 2, Chapter 5 and Chapter 7, 1 August 2008; Chapter 3, 30 April 2002; Chapter 4, 31 July 2003; Part VI: Chapter 5, 28 Sep 11.

2.3-13 "Advanced Series on Ocean Engineering, Volume 15, Random Seas and Design of Maritime Structures", Y. Goda 2000.

2.3-14 "Water Wave Mechanics for Engineers and Scientists", Robert G. Dean and Robert Dalrymple, 1984, Prentice-Hall.

2.3-15 "EurOtop. Wave Overtopping of Sea Defences and Related Structures - Assessment Manual", Eds. Pullen, T., N.W.H AIlsop, T. Bruce, A. Kortenhaus, H. Schuttrumpf &

J.W., van der, Meer, 2007.

2.3-16 "Colorado River - Flood Guide," Lower Colorado River Authority, Texas, January 2003.

2.3-17 "Water for Texas - 2007," Volumes 1,11, and Ill, Texas Water Development Board, January 2007.

Dam Breaches and Failures 2.3-29

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project 2.3-18 WLIDelft, 2005; Delft3D-FLOW, "Simulation of multi-dimensional hydrodynamic flows and transport phenomena, including sediments," published and printed by WLIDelft Hydraulics, Rotterdamseweg, 185, P.O. Box 177, 2600 MH Delft, The Netherlands.

2.3-18a Abbott, M. B., A. Damsgarrd, and G. S. Rodenhuis, System 21, "'Jupiter' (A Design System for Two-Dimensional Nearly Horizontal Flows)", Journal of Hydraulic Research, Vol. 11, No. 1, 1973.

2.3-18b National Weather Service, Dam-Break Flood Forecasting Model, Version 13, User's Manual, January 30, 1982.

2.3-19 South Texas Project Units 1 & 2 Flood Analysis, prepared by Atkins, Austin, Texas for STPEGS, Document No. 120021, March 2012.

2.3-20 Chow, V. T., Open Channel Hydraulics, McGraw-Hill, Inc., 1959.

2.3-21 T. L. Wahl, "Prediction of Embankment Dam Breach Parameters, A Literature Review and Needs Assessment", Dam Safety Research Report DSO-98-004, Dam Safety Office, Water Resources Research Laboratory, U.S. Department of the Interior, Bureau of Reclamation, July 1998.

2.3-22 D. L. Fread and J. M. Lewis, "NWS FLDWAV Model Theoretical Description and User Documentation," Hydrologic Research Laboratory, Office of Hydrology, National Weather Service, U.S. National Oceanic and Atmospheric Agency, Silver Spring, Maryland, 1998.

2.3-23 "User's Guide to RMA2 WES," Version 4.5., Coastal and Hydraulics Laboratory, Waterways Experiment Station, Engineer Research and Development Center, U.S.

Army Corps of Engineers, April 22, 2005.

2.3-24a G. J. Arcement and V. R. Schneider, "Guide for Selecting Manning's Roughness Coefficients for Natural Channels and Flood Plains," Water-Supply Paper 2239, United States Geological Survey, 1989.

2.3-24b "HEC-RAS, River Analysis System, User's Manual," Version 3.1.3, U.S. Army Corps of Engineers, Hydrologic Engineering Center, May 2005.

2.3-25 Fread, D. L., BREACH, "An Erosion Model for Earthen Dam Failures", Hydrologic Research Laboratory, Office of Hydrology, National Weather Service, U.S. National Oceanic and Atmospheric Agency, Silver Spring, Maryland, July, 1988.

Dam Breaches and Failures 2.3-30

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushimna Response Project Table 2.3-1 Summary of the 68 Dams in Colorado River Basin with 5,000 AF or More Storage Capacity No. Dam Name County Height of Length of Top of Dam Maximum Dam Type Date of Dam (ft) Dam (ft) Elevation (ft Capacity (AF at Completion MSL) top of dam) 01 Mansfield Dam Travis 266.4 7,089 750.1 (754.1 3,300,000 [4] Concrete Gravity 1942 ft: top of Earth and Rockfill parapet) 02 Simon Freese Dam [5] Coleman 148 15,950 1584 1,470,000 [4] Earth and Rock Fill 1990 Embankment 03 Twin Buttes Dam [5] Tom Green 134 42,460 1991 1,294,000 [3] Earthfill 1963 04 Buchanan Dam Burnet 145.5 10.987 1025.35 1,180,000 [1] Multiple Concrete 1937 Arch. Gated and Gravity sections 05 Robert Lee Dam [5] Coke 140 21,500 1928 1,074,000 [3] Earthfill 1969 06 0 C Fisher Dam [5] Tom Green 128 40,885 1964 815,000 [2] Earthfill 1952 07 Brownwood Dam [5] Brown 120 1,580 1449.5 448,2000 [1] Earthfill 1933 08 Lake J B Thomas Dam [5] Scurry 105 14,500 2280 431,000 [2] Earthfill 1952 09 Alvin Wirtz Dam Burnet 118.29 5,491 835.25 226,000 [4] Concrete and 1951

_Earthfill 10 Brady Dam [5] McCulloch 104 8.400 1783 213,000 [3] Earthfill 1963 11 Natural Dam [1] [5] Howard 47 [6] [6] 207,265 Earth 1989 12 Tom Miller Dam Travis 85 1,590 519 115,404 [1] Concrete Gravity 1939 13 Coleman Dam [5] Coleman 90 3,200 1740 108,000 [3] Earthfill 1966 14 Champion Creek Dam [5] Mitchell 114 6,800 2109 103,600 [3] Earthfill 1959 15 Cedar Creek Dam Fayette 96 8,000 401 101,000 [4] Earthfill 1977 16 Oak Creek Dam [5] Coke 95 3,800 2104 83,800 [3] Earthfill 1952 17 Colorado City Dam [5] Mitchell 85 4,800 2090 78,400 [4] Earthfill 1949 18 Hords Creek Dam [5] Coleman 91 6,800 1939 66,300 [3] Earthfill 1948 19 Roy Inks Dam Burnet 96.5 1,547.5 922 63,500 [1] Concrete Gravity 1938 20 Mitchell County Dam [1] [5] Mitchell 70 [6] [6] 50,241 Earth 1991 21 Decker Creek Dam Travis 83 6,390 563 45,300 [2] Earthfill 1967 22 Nasworthy Dam [5] Tom Green 50 5,480 1883.5 43,300 [4] Earthfill 1930 Damn Breaches and Failures 2.3-31

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPI & 2 Fukushima Response Project Table 2.3-1 Summary of the 68 Dams in Colorado River Basin with 5,000 AF or More Storage Capacity (Continued)

Dam Name County Height of Length of Top of Dam Maximum Dam Type Date of No. Dam (ft) Dam (ft) Elevation (ft Capacity (AF at top Completion MSL) ofdam) 23 Ballinger Municipal Lake Dam [1] [5] Runnels 76 6,200 1,694 34,353 Earth 1985 24 Elm Creek Dam [1] [5] Runnels 57 5,640 1.810 33,500 Earth 1983 25 Bastrop Dam Bastrop 85 4,000 458 24.200 [1] Earthfill 1964 26 Sulphur Springs Draw Damn [1] [5] Travis 33 [6] [6] 20.692 Earth 1993 27 Upper Pecan Bayou WS SCS Site 7 Callahan 63 3,950 1888.9 20,000 [3] Earthfill 1970 Dam [5]

28 Brady Creek WS SCS Site 17 Dam [I] Mcculloch 50 4,208 [6] 13,511 Earth 1962

[5]

29 Brady Creek WS SCS Site 28 Dam [1] Concho 42 6,459 [6] 13,042 Earth 1957

[5]

30 Brady Creek WS SCS Site 31 Dam [I] Concho 50 5,910 [6] 11,155 Earth 1958

[5]

31 Old Lake Winters City Dam [1] [5] Runnels 37 3,090 1800.2 10,032 Earth 1945 32 Eagle Lake Dam [2] Colorado Varies 6 ft 5,300 Not known 9,600 at EL 170 ft, Earthfill 1990

+/ MSL 33 Brady Creek WS SCS Site 20 Dam [1] Concho 43 4,010 [6] 9,494 Earth 1959

[5]

34 Northwest Laterals WS SCS Site 5A Coleman 57 2,631 [6] 9,416 Earth Dam [1] [5]

35 Max Starcke Dam Burnet 98.8 860 766 [1] 8,760 [1] Concrete with 1951 738 [7] Roof-weir Gated 36 Jim Ned Creek WS SCS Site 25 Dam [1] Coleman 44 2,400 [6] 8,368 Earth 1963

[5]

37 Jim Ned Creek WS SCS Site 12EI Dam Coleman 64 2,000 [6] 8,271 Earth 1965

[1] [5]

38 Ballinger City Lake Dam [1] [5] Runnels 30 4.400 1704.6 8,215 Earth 1947 39 Elm Creek WS NRCS Site 3 Rev. [1] 15] Runnels 39 [6] [6] 8,165 Earth 2004 Darn Breaches and Failures 2.3-32

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STPJ & 2 Fukushimna Response Project Table 2.3-1 Summary of the 68 Dams in Colorado River Basin with 5,000 AF or More Storage Capacity (Continued)

Dam Name County Height of Length of Top of Dam Maximum Dam Type Date of No. Dam (ft) Dam (ft) Elevation (ft Capacity (AF at Completion MSL) top of dam) 40 Clear Creek WS SCS Site 6 Dam [1] [5] Brown 50 2,101 1461 8,083 Earth 1958 41 Jim Ned Creek WS SCS Site 21 Dam [1] [5] Coleman 92 1,915 [6] 7.930 Earth 1963 42 Clear Creek WS SCS Site 4 Dam [1] [5] Brown 45 2,300 1508.6 7,891 Earth 1958 43 Upper Pecan Bayou WS SCS Site 2 Dam [1] [5] Callahan 69 2,025 1948.8 7,833 Earth 1967 44 Brady Creek WS SCS Site 14 Dam [1] [5] Mcculloch 43 4,091 [6] 7.,732 Earth 1956 45 Home Creek WS SCS Site 13 Dam [1] [5] Coleman 45 2,410 [6] 7,679 Earth 1974 46 Valley Creek WS SCS Site 1 Dam [1] [5] Nolan 52 5,100 2121.8 7,600 Earth 1968 47 Upper Pecan Bayou WS SCS Site 24 Dam [1] [5] Coleman 50 1,800 1606.4 7,394 Earth 1972 48 Brownwood Laterals WS SCS Site 3 Dam [1] [5] Brown 83 1.930 1473.9 7.377 Earth 1973 49 Northwest Laterals WS SCS Site 1 Dam [1] [5] Runnels 50 2,520 [6] 7,181 Earth 1964 50 Brady Creek WS SCS Site 32 Dam [1] [5] Concho 32 8,075 [6] 7,053 Earth 1959 51 Longhorn Dam [1] Travis 65 1,240 464 6,850 Earth. 1960

________Gravity 52 Jim Ned Creek WS SCS Site 23 Dam [1] [5] Coleman 62 1,980 [6] 6,754 Earth 1962 53 Elm Creek WS NRCS Site 7 [1] [5] Runnels 39.5 [6] [6] 6,500 Earth 1998 54 Home Creek WS SCS Site 7A Dam [1] [5] Coleman 48 3,396 [6] 6,367 Earth 1970 55 Jim Ned Creek WS SCS Site 12 Dam [1] [5] Coleman 84 1,900 [6] 6,334 Earth 1963 56 Mukewater Creek WS SCS Site 1OA Dam [1] [5] Coleman 35 3,190 1485.7 6,130 Earth 1965 57 Elm Creek Lake Dam [1] [5] Rtnnels 23 450 1635 6,018 Earth 1930 Dain Breaches and Failures 2,3-33

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushimna Response Project Table 2.3-1 Summary of the 68 Dams in Colorado River Basin with 5,000 AF or More Storage Capacity (Continued)

Dam Name County Height of Length of Top of Dam Maximum Dam Type Date of No. Dam (ft) Dam (ft) Elevation (ft Capacity (AF at Completion MSL) top of dam) 58 Clear Creek WS SCS Site 3 Dam [1] Brown 55 1,950 1451.5 5,988 Earth 1960

- _ [5]

59 Se Laterals WS SCS Site 7 Dam [1] San Saba 43 2,225 [6] 5,899 Earth 1968

[5]

60 Brady Creek WS SCS Site 21 Dam [1] Concho 30 3,543 [6] 5,742 Earth 1958

[51 61 Upper Pecan Bayou WS SCS Site 12 Callahan 65 1,400 1759.3 5,707 Earth 1967 Dam [1] [5]

62 Moss Creek Lake Dam [1] [5] Howard 67 2,450 2341.6 5,700 Earth 1939 63 Cummins Creek WS SCS Site 1 Dam Lee 25 4,050 450.9 5,627 Earth 1958

[I]

64 Brady Creek WS SCS Site 36 Dam [1] Concho 33 1,973 [6] 5,352 Earth 1955

[5]

65 Northwest Laterals WS SCS Site 2 Coleman 52 2,082 [6] 5,297 Earth 1964 Dam [1] [5]

66 Jim Ned Creek WS SCS Site 26A Coleman 46 4,000 [6] 5,280 Earth 1966 Dam [1] [5]

67 Jim Ned Creek WS SCS Site 19 Dam Taylor 28 2,985 [6] 5,218 Earth 1960

[1] [5]

68 Clear Creek WS SCS Site 1 Dam [1] Brown 40 1,542 1397.6 5,128 Earth 1960

[5]

[I ] Data provided by TCEQ

[2] Data provided by TWDB: data was directly listed in Reference 2.3-8

[3] Data provided by TWDB: data were extrapolated based on the storage-stage curves in Reference 2.3-8

[4] Data provided by TWDB: data were extrapolated based on the storage-stage area data

[5] Dams located upstream of Buchanan Dam

[6] No information was given by TCEQ

[7] Data from LCRA. in Reference 2.3-16 Dam Breaches and Failures 2.3-34

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Table 2.3-2 500-year and SPF Inflow Peak Discharges at Selected Locations along the Colorado River (in cfs)

Flood Event Buchanan Mansfield Tom Bastrop Garwood Wharton Bay City Miller 500-year 382,400 499.700 366,900 321,900 256.700 204.700 187,900 SPF 484,800 737,000 402,500 359,900 285,500 237,800 214,200 Source: Reference 2.3-10 Table 2.3-3 Breach Parameters for Buchanan and Mansfield Dams Breach Parameters Buchanan Dam Mansfield Dam Average Width of Breach (ft) 1470 1360 Breach Bottom Elevation (ft, MSL) 879.8 484 Breach Top Elevation (fi, MSL) 1,028.4 757 Side Slope of Breach 0 0 Breach Time to Failure (hrs) 0.1 0.1 Dam Breaches and Failures 2.3-35

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 FukushimaResponse Project Table 2.3-4 Initial Estimation of Manning's Roughness Coefficient n Values Assigned to the USGS NLCD Dataset USGS Classification Grid-Code Description n Value II Open water 0.03 21 Low intensity residential 0.07 22 High intensity residential 0.09 23 Commercial/indtstrial/transportation 0.10 31 Bare rock/sand/clay 0.04 32 Quarries/strip mines/gravel pits 0.035 41 Deciduous forest 0.095 42 Evergreen forest 0.085 51 Shrubland 0.08 71 Grasslands/herbaceous 0.04 81 Pasture/hay 0.045 82 Row crops 0.05 83 Small grains 0.055 85 Urban/recreation grasses 0.03 91 Woody wetlands 0.10 92 Emergent herbaceous wetlands 0.085 Source: Reference 2.3-10 Dant Breaches and Failures 2.3-36

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STPJ & 2 Fukushima Response Project Table 2.3-5 Estimated Water Levels due to Dam Break, Wind Setup, and Wave Run-up at STP I & 2 Power Block Structures Dam Break Wind Setup Wave Run-up Water Level Fetch Water Level (ft) (ft) (ft MSL)

(ft MSL)

A 28.6 2.9 4.7 36.2 B 28.6 3.0 4.5 36.1 Table 2.3-6 Estimated Water Levels due to Dam Break, Wave Transmission and Wave Run-up at ECW Intake Structure ECP Crest Elevation Wave Run- Water Level at STP Site (ft MSL) up (ft) (ft MSL) 34 1.8 35.8 Dam Breaches and Failures 2.3-37

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushinia Response Project Table 2.3-7 Ground, Maximum Water and Buoyancy Elevations from the MCR Breach Analysis of the UFSAR for STP I & 2 (Reference 2.3-1)

Reactor Mechanical- Fuel- Diesel- ECW Auxiliary Contaunmen Electrical Handling Generator Intake Feecdvater Building Auxiliaries Building Building Structure Storage Tank Building N S E W N S E W N S E W N S E W N S E W N S E W Ground Elevation 28.0 28.0 28.0 28.0 28.0 28.0 28.0 28.0 28.0 28.0 28.0 28.0 25.0 25.0 10.0 34.0 28.0 28.0 28.0 28.0 Max. Water Elevation 49.0 44.5 50.6 48.1 50.6 50.8 50.6 50.3 44.5 44.5 44.5 44.5 40.8 40.8 40.8 40.8 50.0 50.0 50.0 50.0 Buoyancy Elevation 49.0 44.5 50.6 48.1 50.6 50.8 50.6 50.3 44.5 44.5 44.5 44.5 40.8 40.8 40.8 40.8 50.0 50.0 50.0 50.0 Table 2.3-8. Maximum Velocities, Depths, and Water Elevations for Selected Locations from the 2012 MCR Breach Analysis (Reference 2.3-19)

Dam Breaches and Failures 2.3-38

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushinra Response Project Table 2.3-9 Debris Characterization Table from the 2012 MCR Breach Analysis (Reference 2.3-19)

DerisItemSzh ra Specific fcto Buoa- nua_

Deis item Size Weit GraV.t Coefficient Ceffice area

_ 1F et ( U )~ fsq ft ) ( Feet)

Gay particles small "Embankmem 0,1-1.0 1.T7 0.27 0.05 Material G~ayBalls Soil Cement Pieces 0-5-3.0 2.37 1.05 1.5 Single cab pick-up 20 5000 0.1 0.1 36 24 34 3000 1.05 16 Vehicles 4-1Wsedan Mobile office traier 30 10000 1.05 0: 1 28 2 Dry WalI pieces 7 30 Office Cublde wal~s 5 60 Contents ChaIrs 2 20 0L7 0.1 2 S'xS' Tables 3 30 0.7 0.1 2 Paper I Sheet metal from buldL's 8 200 Miscellaneous Coth White Rock found In PA 0.125-0.175 2.5 0.4 0.05 IMisc Garbage 0.3 1 1 Dam Breaches and Failures 2.3-39

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Table 2.3-10 Proportions of Debris Items Predicted to be Retained in the ECP from the 2012 MCR Breach Analysis (Reference 2.3-19)

Debris Item Unit I Unit 2 NSC Clay particles Embankment Clay Balls Material Soil Cement Pieces 0.57 0.35 0.68 0.00 0.00 0.00 Single cab pick-up 0.00 0.00 0.00 Vehicles 4-Dr sedan 0.00 0.00 0.00 Mobile office trailer 0.00 0.00 0.00 Dry Wall pieces 0.00 0.00 0.00 Cubicle walls 0.00 0.00 0.00 Office Contents Chairs 5 by 5-ft 1.00 1.00 0.00 Tables 1.00 1.00 0.00 Paper Sheet metal from buildings 0.00 0.00 0.00 Cloth 0 0 0 Miscellaneous White Rock found in power block . . 0 Miscellaneous Garbage 1.00 1.00 0.56 Dam Breaches and Failures 2.3-40

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Table 2.3-11 Extrapolated Inundation Period at Selected Locations from the 2012 MCR Breach Analysis (Reference 2.3-19)

Breach Inundation Time (hours) till 0.25-ft Depth at Selected Locations Scenarios Point 1 Point 2 Point 3 Point 4 Point 5 Point 6 Point 7 Point 8 Point 9 Point 10 Unit 1 61.54 55.21 53.60 57.05 58.41 56.25 30.36 40.26 55.73 64.43 Unit 2 54.69 63.96 54.65 64.43 58.41 57.83 30.65 38.43 56.39 65.86 NSC 51.19 53.53 46.21 53.07 47.83 44.70 21.82 39.10 44.65 53.07 Dam Breaches and Failures 2.3-41

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Dam Location Key I MANSIELD DAM 2 S RA FREESE DAM 3 w4I BUTTESEDAN 4 JCMi94ANj DAM R OSERT LEE DAM 6a CFIUER DAM 7 L" AKEBMV4WOO DAM A JTHOM",S S DAM 9 =AL WIRTZDAM 10 RADYDAM 11 -NATURAL DAM OMMILER DAM 3 MANDAM 09 1 14 A&UCN QEK DAM Is DARCREEKDAM s K CREEKDAM 1?

7 E COLORAOO CITYDAM 19R DAM ITCHE.LCOUNTYRESERVOIR DAM 21 D ECKERCREEKDAM 22 SWORTHY DAM 235mUE4ER tARXICIAL LAKEDAM 25t AKE&ASTfROP DAM 2SU PPEfR 28 SPLINGS ULJLPHLUR PECAN BAYOUIWS SDAM DRAW SITE7 DA 2 ADYCREE WS SUS SE 17 DAM 30BSR GAIWS SS SUrE29DAM YE 31S ADYCRE(E V SCS SITE 31DAM 31 0 LAKEV4TERS CITYDAM Figure 2.3-1a Locations of Dams with Storage Capacity Over 10,000 AF in the Colorado River Basin Upstream of the STP Site Dam Breaches and Failures 2.3-42

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Dam Location Key W AGLELAKEDAM 33 RAYCREEK WS SCS SE 20 DAM 34 N.rm tWESTr LATERALS WS =9S SITE 5A DAM 20 .STARCKE DAM 37 M NED CREEK WIS CSSITE25 DAM 374 NMNEDCREEKWS CS SITE12EI DAM B9 8 34ER CMl LANEDAM SLMCREEK INSNRCS SrTE3 REV 40 CLEARCREEKSISCS SE 8 DAM 41 j IMNEDCREEK WI SOSS*TE21 DAM 42 CLEARGREEKWS CE S(TE4 DAM 43UP PERPECANBAYOUWS SOSBTE 2 DAM 44a RACYCREEK WS SCSSfTE 14DAM 45H OMECREEK WS SCS S`TE 13DAM 46 ALLEYCREEKWS S9S S9TEI DAM 4V PERtMAN BAYOUWS SCS SITE24 DAM 40 "lRDW9WOODLATERALS W SOSSITE 3 DAM 4, INVESTLATERALS WE SOS STE I DAM w0 -"956YCREEK WS SmsE 32 DA*

518-(IHO*O LONGSVRN DAM 524 IMNED CREEK WI SOSSTE 23 DAM LMCREEK WISNRCS FTE7 REV SOE CREEK WS CS STE 7ADAM J NIMNED CREEKWI SCS9STE12 DAM IUJEM R CREEKWS SCS SITE1GADAM 5 7 M LMCREEK LAKEDAM Se ECLEARCREEKIWE SOSSITE3 DAM 5 E LATERALS r WE SOSSITE DAM T B6RA.Y CREEKWSSOS STE 21 DAM 91 UPPER PECANBAYOUJ WE SCSSITE12 DAM 92 DESCREEKLAKEDAM 83'SMNAwI CREEKIN SOSSITEI DAM 64 5 RACYCREKWS SOSSITE30DAM as IWEST LATERALS INS SOSSITE2 DAM N NSIV6D CREEKWNS SOSSITE26AOM b7J IMNED CREEK WS SCSSITE19 DAM 6s JCLEAR CREEK WESSS SITEI DAM Figure 2.3-1b Locations of Dams with Storage Capacity of 5,000 AF to 10,000 AF in the Colorado River Basin Upstream of the STP Site Dam Breaches and Failures 2.3-43

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project llfflr-I Legend

/

Ground Bank Sta 1050 Final Breach 1000 Iv 950 900 Or~n 0 2000 4000 60 8000 10000 12000 14000 Stabon (ft)

Figure 2.3-2 Model Cross-Section at Buchanan Dam w 650.

_w 600, 1000 2000 3000 4000 5000 Station (ft)

Figure 2.3-3 Model Cross-Section at Mansfield Dam Dam Breaches and Failures 2.3-44

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project o 5b 9 1 21 3 4 FI Figure 2.3-4 Locations of Model Cross-sections in the Dam Break Analysis Dam Breaches and Failures 2.3-45

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Legend Ground Bank Sta

.2r (U

0 500 1000 1500 2000 2500 3000 Station (ft)

Note: Between Buchanan and Mansfield Darns and about 49.6 River Miles Upstream of Mansfield Darn.

Figure 2.3-5 Model River Cross-section at About 365 River Miles Upstream of the GIWW Dam Breaches and Failures 2.3-46

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Legend Ground Bank Sta 0) ul 0 2000 4000 6000 8000 10000 12000 14000 16000 18000 Station (ft)

Note: Downstream of Mansfield Dam and about 153 miles Upstream of STP Site.

Figure 2.3-6 Model River Cross-section at About 163.5 River Miles Upstream of the GIWW Danm Breaches and Failures 2.3-47

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Legend Ground Bank Sta w

-100000 -80000 -60000 -40000 -20000 0 20000 Station (ft)

Note: Near the STP site.

Figure 2.3-7 Model River Cross-section at About 10.5 River Miles Upstream of the GIWW Dam Breaches and Failures 2.3-48

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Legend Ground Bank Sta

._o 0

-120000 -100000 -80000 -60000 -40000 -20000 0 20000 Station (ft)

Figure 2.3-8 Model River Cross-section at Downstream Model Boundary at about 0.9 River Miles Upstream of the GIWW Dain Breaches and Failures 2.3-49

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Plan: One n20 Rlver Colorado Reach: I RS: 89146.04 I

07 12 17 22 27 01 Jan2000 01Feb2000 Time Note: Vertical Datum is NAVD 88; model start date was selected arbitrarily.

Figure 2.3-9 Based Case Flood and Stage Hydrographs at the STP Site Dam Breaches and Failures 2.3-50

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Plarr Dam Break nO Rew. Colorado Reach: 1 RS: 89146.04 Jan2000 O0Feb2000 Time Note: Vertical Datum is NAVD 88; model start date was selected arbitrarily.

Figure 2.3-10 Sensitivity Case Flood and Stage Hydrographs at the STP Site Dam Breaches and Failures 2.3-51

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STPJI & 2 Fukushima Response Project 1200-II I Legend WS Max WS 1000-Cris Max W S Ground I

I

-200 1000000 1500000 2'000000 2500000 Masi Ctiannel Distance (ft)

Note: Vertical Datum in NAVD 88.

Figure 2.3-11 Base Case Simulated Maximum Dam Break Surface Profiles from Buchanan Dam to 4,600 ft upstream of GIWW (Vertical Datum in NAVD 88)

Dam Breaches and Failures 23-52

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STPJI & 2 Fukushima Response Project ai A i 1200-1000-I F-Le~

WS maiWS Crit Max WS 800- I Ground S 600-400-ZI 200

-200 0500000 -0 1000000 1500000 2006000 2506000 Msin Channel Distance (ft)

Note: Vertical Datum in NAVD 88.

Figure 2.3-12 Sensitivity Case Simulated Maximum Dam Break Surface Profiles from Buchanan Dam to 4600 ft Upstream of GIWW Dam Breaches and Failures 2.3-53

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project

/ 7 j ~

/

/ SAN BERNARD RIVER BASIN fl*ATOJ RIVER BASIN COLORADO

-..- 1 EL C"P0 N.

N -

/040

--- WATEFr.0E BOUNDARY UNES *94$/4, SCALE Figure 2.3-13 Fetch Directions and Length Dam Breaches and Failures 2.3-54

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-14a ECP and ECW Intake Structure for STP I & 2 Dam Breaches and Failures 2.3-55

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project Figure 2.3-14b ECP Embankment Cross-section Figure 2.3-15 Wave Transmission on Sloped Structure (Reference 2.3-12)

Dam Breaches and Failures 2.3-56

Enclosure NOC-AE-13002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project NO 2 ICO..3S1)

MODEL NOI

-0IE NO2-KI SOUTH TEXAS PROJECT UNITS 1 & 2 SCALE LIM17S ANDO Or COARSE ANALYSIS T*O DIMENSIONAL F-Q, 2.44.7 Revision 0 Figure 2.3-16 Model Configurations from the UFSAR for STP 1 & 2 (Reference 2.3-1)

Dam Breaches and Failures 2.3-57

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 FukushirnaResponse Project so I - - -1 2

a:

35 -

w IC-I--UN NO.2 V,

02 ?0 NIo. I I-.

42 I i LIMITS OF I-FNrGRID

/ MODE9

/ I fa 155 Ii COLUMN NUMSERS 0 '000 2000 3000 rET SCALE,1 NOTE:

ONLIf EVERY FIFTH GRi:;LINE' FOR CLARITY OF PIRESENTATIMN

~S ~NEHOWN GR.D SOUTH TEXAS PROJECT SPACE EQUALS 210 FEET EACH021RECT1ON. UNITS* &2 COARSE GRID AND PLANT LAYOUT FOR MODEL MO.2 AIAALVI.Y1 FIgure 2.4.4-11 Revision 0 Figure 2.3-17 Coarse Grid and Plant Layout for Model No. 2 (Reference 2.3-1)

Dam Breaches and Failures 2.3-58

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project 0 SO 0:0 "y6ZLC . F11' SOUTH TEXAS PROJECT NOTE ONLY EVERY FIFTH GRIDOLIME IS SOWN UNITS 1 & 2 FOR CLARIMY OF PREICNTATION, OME GRID SPACE EQUALS 70 FEET EACN 0RECTrIoN.

FINEGRIDAND "LANTLAYOUI FOR MODEL NO 2 ANALYSIS

gle L4,.*1-2 RCev.140 0

Figure 2.3-18 Fine Grid and Plant Layout for Model No. 2 (Reference 2.3-1)

Dam Breaches and Failures 2.3-59

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project

_400 SCALE IN FEET SOUTH TEXAS PROJECT UNITS 1 & 2 GRID POSITIONS ON PLANT LAYOUT FOR FINEGRJD ANALYSIS Figura 24.4-13 Rev.$1on 0 Figure 2.3-19 Fine Grid at STP Units 1 & 2 for Model No. 2 (Reference 2.3-1)

Damn Breaches and Failures 2.3-60

-I

'50 (D

0 NOTE' EhVELOPES OF MAXIMUiM WATER SURFACE ELEVATIONS SOUTH TEXAS PROJECT Q) ALONG WECT SIDE OF UNIT NO.I. 1890' BREACH UNITS 1 & 2 6D A:.ONg EAST SIDE OF UNIT NO.Z, 1890' BREACH ALONG WEST SIDE OF UNIT NO.I. 420' BREACH ALONG EAST SIDE OF UNIT NO.2, 420' BREACH PROFTES SHOWN ARE BASED ON ENVELOPE OF MAXIMU* WATER SURFACE ELEVATIONS RESULTING FLOW PROFILES BETWEEN STRUCTURES FROM CALCULAT'ONS UTILIZING MODEL NO 2 420 FT. BREACIhVS. 1890 FT. BREACH AND A MANNINOS n VALUE OF 0.046 Sz Figure 2.4.4.9 Revision 0

~. >

Sm i500 (31 CD

P.O Am 0

0 0

-4a "aSC ns 90

-u "a)C

-I

-u C

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project 2,200 m 1,810 m 2,980 m 20x 20m gridsize 10 x 20mgrid size 20 x 20mgrid size E

N 20 x10m g~d size 10 xl10mgrid size 20 xl10mgrid size STP E 1&12 E,

t 1 20 x20 siz 20 mgrid size 20x 20m gridsize I I

/ 6,990 m (22,933.1 It)

E 10,000 m (32,808.4 It)

N 12,900 m (42,322.8 It)

Figure 2.3-22 Layout of Modeling Domain for the Independent MCR Breach Analysis with DeIft3D Dam Breaches and Failures 2.3-63

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project

-- FLDWAV - BREACH 140,000 120,000 100,000 80,000 60,000 40,000 20,000 0

0 246 810 12 Time (hours)

Figure 2.3-23 Flood Hydrograph From FLDWAV and BREACH Simulations Dam Breaches and Failures 2.3-64

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-24 Layout of Grid Model for STP 3 & 4 COLA (with breach facing STP 3) (Elevation datum is NGVD 29)

Dam Breaches and Failures 2.3-65

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-25 Land Cover Types Assigned (STP 3 & 4 COLA)

Dam Breaches and Failures 2.3-66

NOC-AE-13o029 75 Flooding Hazard Reevaluatiot RePort Fukushima Response Project Figure 2.3-26 Land cover Types Asigned (Reference 2.3-I9)

Dam Breaches and Failures 2.3-67

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-27a Unit I Breaching Scenario Water Levels (Reference 2.3-19)

Dam Breaches and Failures 2.3-68

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPJ & 2 Fukushima Response Project Figure 2.3-27b Unit 2 Breaching Scenario (Reference 2.3-19)

Dam Breaches and Failures 2.3-69

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-27c NSC Breaching Scenario (Reference 2.3-19)

Dam Breaches and Failures 2.3-70

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-28 Location of Monitoring Points (Reference 2.3-19)

Dam Breaches and Failures 2.3-71

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project Figure 2.3-29a Unit I Breach Scenario Maximum Velocity Vectors (Reference 2.3-19)

Dam Breaches and Failures 2.3-72

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project Figure 2.3-29b Unit 2 Breach Scenario Maximum Velocity Vectors (Reference 2.3-19)

Dam Breaches and Failures 2.3-73

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STPJ & 2 Fukushima Response Project Figure 2.3-29c NSC Breach Scenario Maximum Velocity Vectors (Reference 2.3-19)

Dam Breaches and Failures 2.3-74

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Unit 1 Breach Results - Velocities at Selected Locations 14 12 PointI Point2 10 Point 3

- - - Point 4 I. 8 Point 5 6 -Point 6

... Point 7 4 --- Points

-- Point 9 2

Point 10 0

0 5 10 15 20 25 30 35 Ti* (hoM")

Figure 2.3-30a Unit I Breach Scenario Velocity Time History (Reference 2.3-19)

Dam Breaches and Failures 2.3-75

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 FukushimaResponse Project Unit 2 Breach Results - Velocities at Selected Locations 12 Point 1 10 Point 2

- Point 3 8

-- - Point 4 aI Point 5 6

-Point6 4 Point 7

- - - Point 8 2 --- Point9 Point 10 0

0 5 10 15 20 25 30 35 Time (hows)

Figure 2.3-30b Unit 2 Breach Scenario Velocity Time History (Reference 2.3-19)

Dam Breaches and Failures 2.3-76

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project NSC Breach Results - Velocities at Selected Locations 10 - -.----------- ---.-.-.-..--

PointI PoInt 2

- Point 3

--- Point 4

- Point 5

Point 6 6 Pi Point 7

-- -- - -- -- PointS

.. Point9 0 " ' -- -------- -- P in

.- Point 10 0 5 10 15 20 25 30 35 Ti" 2onit Figure 2.3-30c NSC Breach Scenario Velocity Time History (Reference 2.3-19)

Dam Breaches and Failures 2.3-77

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushimna Response Project 2.4 Storm Surge This section addresses potential flooding at the STP 1 & 2 safety-related structures, systems and components (SSCs) due to the probable maximum storm surge (PMSS). The STP 1 & 2 safety-related SSCs include those located in the power block area and those in the Essential Cooling Pond (ECP) including the Essential Cooling Water Intake Structure (ECWIS), cooling water discharge structure and south ECP embankment (see Figures 1.1-3a through 1.1-3e).

The current design basis (CDB) flood elevations of the safety-related SSCs at STP 1 & 2 are governed by the maximum flood levels resulting from the postulated MCR embankment breach event, as documented in Table 2.4.4-3 of the UFSAR (Reference 2.4-1). The CDB flood elevations in the power block vary from a minimum of 44.5 ft MSL at the Diesel Generator Building and the north face of Mechanical Electrical Auxiliaries Building to a maximum of 50.8 ft MSL at the south face of the Fuel Handling Building. In the ECP, the CDB flood elevation was established to be 40.8 ft MSL at the ECWIS Section 1.2 of this report provides further details of the CDB flood elevations on various safety-related SSCs.

PMSS flooding impact was evaluated for the STP 1 & 2 site as described in the STP 1 & 2 UFSAR (Subsection 2.4.5.2, Reference 2.4-1) that used a combination of a one-dimensional bathystrophic model to simulate the storm surge offshore of the Colorado River mouth and a one-dimensional hydraulic model HEC-2 to simulate the routing of the flood surge up the Colorado River. The resulting PMSS still water level at the site was estimated at 26.78 ft MSL, below the power block grade elevation of 28 ft MSL. Consequently, UFSAR concluded in Subsection 2.4.5.3 that consideration of coincidental wind-wave actions was not warranted. The PMSS still water level of 26.78 ft MSL was also bounded by the design basis flood elevations from the postulated MCR embankment breach.

The STP 3 & 4 Combined License Application (COLA) (Reference 2.4-2) also examines the PMSS flooding impacts at the site. The COLA (Subsection 2.4S.5) uses three different methods including the combination of one-dimensional bathystrophic and hydraulic model as in STP 1 &

2 UFSAR, extrapolation of surge results from Categories I to V storms from the two-dimensional storm surge model SLOSH (as compiled by the National Oceanic and Atmospheric Administration, NOAA) and simulation by the state-of-the-art storm surge model ADCIRC.

This reevaluation effort includes a comprehensive review of the UFSAR and COLA storm surge model studies to determine if the approach and methodology, modeling tools, supporting data and results meet the present day requirements specified in NRC 50.54(f) Request for Information (RFI) letter of March 12, 2012. In addition to the numerical modeling software used, input data such as model grid, initial and boundary conditions, conformance with regulatory guidance and industry standards, specifically the NRC Interim Staff Guidance (ISG) JLD-ISG-2012-06 (Reference 2.4-3), are reviewed.

Review of the previous storm surge evaluations, supplemented with other more recent storm surge analyses, indicates that the one-dimensional bathystrophic storm surge and hydraulic routing modeling approach adopted in the UFSAR does not meet the present-day requirements of the 50.54(f) RFI letter. The SLOSH model extrapolation approach presented in the STP 3 & 4 COLA is also not considered to be a present-day methodology. Accordingly, the storm surge model simulation results from the two-dimensional ADCIRC model is used as the basis for the PMSS reevaluation for the STP 1 & 2 site. The reevaluation also follows the guidance of ANSI/ANS 2.8-1992 (Reference 2.4-4) on the PMSS assessment and the specification of the antecedent and combined event conditions, consistent with the recommendations of Storm Surge 2.4-1

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushinta Response Project NUREG/CR-7046 (Reference 2.4-5). The reevaluation demonstrates that the PMSS and coincidental wave actions do not constitute the current flood design basis, which is controlled by the MCR breach flooding event providing the bounding flood levels for the plant.

2.4.1 Probable Maximum Winds and Associated Meteorological Parameters The hydrometeorological conditions that would produce the probable maximum meteorological wind (PMMW) at the STP 1 & 2 site would be due to the PMH. The PMH is described in the National Oceanic and Atmospheric Administration (NOAA) National Weather Service Technical Report No. 23 (NWS 23, Reference 2.4-6, p. 2) as "a hypothetical steady state hurricane having a combination of values of meteorological parameters that will give the highest sustained wind speed that can probably occur at a specified coastal location." The meteorological parameters that give the highest sustained wind speed for the PMH windfield include the peripheral pressure (Pn), central pressure (p,), radius of maximum winds (R), forward speed (T), track direction (8), and inflow angle (4)).

The initial PMH parameters are obtained based on the milepost location along the Gulf Coast (Reference 2.4-6, p. 4). The PMH parameters in NWS 23 were established based on data from historical hurricanes between 1851 and 1977.

The PMH parameters as determined from NWS-23 (Reference 2.4-6) are listed in Table 2.4-1.

The peripheral pressure is 30.12 in Hg. The central pressure is 26.19 in Hg. Therefore, the PMH central pressure deficit (AP) is estimated to be 3.93 in Hg, or 133.0 millibars. The radius of maximum winds had upper and lower limits of 5 and 21 nautical miles, respectively. The forward speed had upper and lower limits of 6 and 20 knots, respectively.

The effect of long-term climate variability on hurricane intensity is an area of active research.

Since 1977, several intense hurricanes had made landfall on the Gulf of Mexico and Atlantic coasts. Research on the effects of El Niiio/Southern Oscillation indicated that while El Niho conditions tend to suppress hurricane formation in the Atlantic basin, La Nih~a conditions tend to favor hurricane development (Reference 2.4-7). Additionally, research indicated possible relationship between the Atlantic Multi-decadal Oscillation (AMO) and hurricane intensity (Reference 2.4-7). AMO is defined as the variation of long-duration sea surface temperature in the northern Atlantic Ocean with cool and warm phases that may last for 20 to 40 years. It shows that hurricane activities increase during the warm phases of the AMO compared to hurricane activities during the AMO cool phases. Recent hurricane data indicate that Atlantic hurricane seasons have been significantly more active since 1995. However, hurricane activities during the earlier years, such as from 1945 to 1970, were apparently as active as in the recent decade (References 2,4-7 and 2.4-8).

Blake et al. (Reference 2.4-8) indicated that during the past 40 years, the conterminous U.S.

was affected by the landfall of three Category 4 or stronger hurricanes: Hurricane Charley of 2004, Hurricane Andrew of 1992, and Hurricane Hugo of 1989. Based on the analysis of hurricane data from 1851 to 2010, they summarized that on average the U.S. is affected by a Category 4 or stronger hurricane approximately once every 8 years, thereby suggesting that there have been fewer exceptionally strong hurricane landfalls during the past 40 years than an expected 40-year average of approximately five (Reference 2.4-8).

Because NWS-23 (Reference 2.4-6) includes the last active hurricane period from 1945 to 1970 (and any such earlier periods from 1851) in the analysis, it is reasonable to assume that the Storm Surge 2.4-2

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project PMH parameters derived are sufficiently conservative even in the considerations of future climate variability.

2.4.2 Storm Surge Water Levels The storm surge is defined as "an abnormal rise of water generated by a storm, over and above the predicted astronomical tide" (Reference 2.4-9, p. 1). The storm surge coinciding with a hurricane typically lasts several hours and affects about one hundred miles of coastline (Reference 2.4-9, p. 1). The setup of the storm surge from the hurricane occurs due to the action of surface wind stress and due to atmospheric pressure reduction. Generally, the storm surge is taken as the sum of several components, including sea level initial rise, astronomical tides, wind setup, setup due to atmospheric pressure reduction, and setup due to breaking waves (Reference 2.4-10, p. 20).

The STP 1 & 2 site is located adjacent to the Lower Colorado River approximately 16 miles upstream from the Gulf of Mexico in Matagorda County, Texas (Figure 1.1-1). The natural ground at the site varies in elevation from approximately 25 ft to about 30 ft MSL. On the south and southeast, the STP 1 & 2 site is fringed by the plant access road (East Access Road),

Texas FM 521 and Service Road. The PMSS at the site would inundate the vast area from the Gulf of Mexico shoreline before affecting the STP 1 & 2 site.

2.4.2.1 Historic Storm Surge Events A list of hurricanes that have impacted the Texas Coast from 1900 to 2005 is shown in Table 2.4-2. Figure 2.4-1 and Figure 2.4-2 depict hurricane tracks that have impacted the Texas Coast from 1852 to 2006 (Reference 2.4-11). A frequency analysis of hurricanes occurring between 1900 and 1963 along the Gulf Coast of Texas noted that "dangerous and destructive tropical cyclones (hurricanes) can be expected to cross the Texas Coast on the average of about once every three years" (Reference 2.4-12, p. 1). Table 2.4-2 indicates the frequency of hurricanes impacting the Texas Coast between 1900 and 2005 is still about once every three years. Blake et al. (Reference 2.4-8) compiled and reported recent hurricane data between 1851 and 2010.

They found that the Texas coast has experienced approximately 64 (Category 1 and above) hurricane landfalls during this 160-year period. After 2005, two hurricanes made landfall on the Texas coast, the Category 1 (in Saffir-Simpson Hurricane Scale) Hurricane Dolly and the Category 2 Hurricane Ike, both in 2008 (Reference 2.4-8).

As the Texas coast has a relatively gentle land slope with low-lying coastal elevations, the storm surge resulting from these hurricanes is capable of flooding significant land areas. For example, Reference 2.4-13 (p. 40) states that "reported surges were 16.6 feet MSL at Port Lavaca, 14.5 feet MSL at Port 0' Connor, 15.2 feet MSL at Matagorda, and 14.8 feet MSL on the upper Houston Ship Channel. A high water line varying from 15.7 to 22 feet MSL, established from debris near the head of Lavaca Bay, probably included the undetermined effects of wave setup and runup." The peak storm surge elevation near STP was about 16 feet MSL from Hurricane Carla in 1961 (Reference 2.4-13, p. 46).

2.4.2.2 Storm Surge Analysis The numerical simulation model "Advanced Circulation (ADCIRC)", which is a hydrodynamic circulation code that simulates water level and current over an unstructured gridded domain, was used to simulate the PMSS elevation at the STP site in support of the Combined License Application (COLA) for STP 3 & 4 (Reference 2.4-2, FSAR Subsection 2.4S.5, Rev. 07).

Storm Surge 24-3

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project ADCIRC can be run as a two-dimensional (2D) or three-dimensional (3D) model and is used for modeling tidally driven and wind and wave driven circulation in coastal waters; forecasting hurricane storm surge and flooding; inlet sediment transport/morphology change studies, and dredging/material disposal studies.

The ADCIRC model, which has been developed over the past 20 years, was selected in recognition of the fact that "current best practices" for predicting storm surge are evolving rapidly due to the very high level of interest and active involvement of the Federal Emergency Management Agency (FEMA), the NOAA, and the U. S. Army Corps of Engineers (USACE).

Associated supporting research has been ongoing at several major universities. These ongoing efforts have resulted in major improvements to this more complex multidimensional computer model used to predict storm surge. Additionally, digital elevation maps based on Light Detection and Ranging (LIDAR) for use with ADCIRC were recently made available for a wider area, including the STP site. The LIDAR based maps improve the accuracy and resolution of the topographic grid, an important input to the computer models, such as ADCIRC, that predict storm surge. Assumptions and initial conditions used with the ADCIRC model are described below.

The ADCIRC two-dimensional depth-integrated model (ADCIRC-2DDI, version 49) was used to perform the hydrodynamic computations to estimate storm surge levels at the STP site. This model uses depth-integrated equations of mass and momentum conservation subject to incompressibility, Boussinesq, and hydrostatic pressure approximations (References 2.4-14 through 2.4-17). The water elevation is obtained from the solution of a depth-integrated continuity equation in Generalized Wave-Continuity Equation (GWCE) form. Velocity is obtained from the solution of the 2D momentum equations. All nonlinear terms have been retained in these equations. ADCIRC is run using either a Cartesian or a spherical coordinate system. The GWCE can be solved using either a consistent or a lumped mass matrix, and an implicit or explicit time stepping scheme.

ADCIRC is linked to a computer program called SWAN that calculates the wave-induced setup in addition to the wind-induced setup calculated by ADCIRC. SWAN is a third-generation wave model developed by the Delft University of Technology. SWAN computes random, short-crested wind-generated waves in coastal regions and inland waters (Reference 2.4-18). The unstructured-mesh SWAN spectral wave model and the ADCIRC shallow-water circulation model have been integrated into a tightly coupled SWAN + ADCIRC model. Hurricane waves and storm surge as estimated by the coupled SWAN + ADCIRC model have been validated for Hurricane Katrina and Hurricane Rita, demonstrating the importance of inclusion of the wave-circulation interactions.

ADCIRC boundary conditions include specified elevation (harmonic tidal constituents or time series), specified normal flow (harmonic tidal constituents or time series), zero normal flow, slip or no slip conditions for velocity, external barrier overflow out of the domain, internal barrier overflow between sections of the domain, surface stress (wind and/or wave radiation stress),

atmospheric pressure, and outward radiation of waves (Sommerfield condition).

FEMA certified ADCIRC for use in performing storm surge analyses as part of their program for developing Flood Insurance Rate Maps (FIRMs) along coastal areas of the United States. This model is the standard coastal model used by the USACE. In addition to USACE projects, ADCIRC is used by the National Oceanic and Atmospheric Administration (NOAA) and the Naval Research Laboratory (NRL).

Storm Surge 2.4-4

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project The antecedent water level, as defined in RG 1.59 (Reference 2.4-19), is estimated separately and used to establish the model initial water level. The PMH parameters (Ap, radius of maximum wind, forward speed, track direction), as described in Subsection 2.4.1, are used to define the physical attributes of the PMH in the model. Model simulations are performed with numerous combinations of input PMH parameters to obtain the PMSS elevation. The effect of wind-wave run-up is superimposed on the PMSS elevation to obtain the maximum water level at the STP facilities.

2.4.2.2.1 Antecedent Water Level According to RG 1.59 (Reference 2.4-19), the 10 percent exceedance high spring tide including initial rise should be used to represent the maximum storm surge antecedent water level. RG 1.59 defines the 10 percent exceedance high spring tide as the high tide level that is equaled or exceeded by 10 percent of the maximum monthly tides over a continuous 21-year period. For locations where the 10 percent exceedance high spring tide is estimated from observed tide data, RG 1.59 indicates that a separate estimate of initial rise (or sea level anomaly) is not necessary.

For the STP site, the 10 percent exceedance high spring tide was estimated from the tidal records at the NOAA Freeport, Texas tide gage station, the closest tidal station from STP which has 21 years of data. The Freeport station is located approximately 45 miles southeast of the site. The daily high/low tide elevations for the period from January 1987 to December 2007 are used for the 10 percent exceedance high spring tide estimate. The data period include missing data from October 1994 to March 1995, which were populated by the maximum monthly tidal levels for the corresponding months between 1961 and 1986. The missing data period was not affected by any hurricane or tropical storm in the Gulf of Mexico. The 10 percent exceedance high spring tide elevation at this station was found to be 3.59 ft NGVD 29 (National Geodetic Vertical Datum of 1929, which is the same as MSL). Because the 10 percent exceedance high spring tide is estimated from tidal records, as recommended by RG 1.59, no additional assessment for initial rise was performed.

In addition to the 10 percent exceedance high spring tide and initial rise, the long-term trend observed in tide gage measurements is also considered to account for the expected sea level rise for the 100-year period. The long-term sea level rise trend at Freeport, Texas, as estimated based on data from 1954 to 2006, is 1.43 ft per century. Accordingly, a nominal long-term sea level adjustment was applied to the 10 percent exceedance high tide level resulting in an antecedent water level of 5.1 feet NGVD 29. This water level was converted to approximately 4.9 ft NAVD 88 (North American Vertical Datum of 1988) and was used as the initial water level in the ADCIRC model simulations.

2.4.2.2.2 STP 3 & 4 COLA ADCIRC Model Grid Topographic and bathymetric data for the model domain are essential to correctly simulate offshore and inland hurricane surge propagation. In addition to describing bathymetry and topography, the model grid should account for pronounced vertical features that have small horizontal scales relative to the grid scale.

Storm Surge 2.4-5

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 2.4.2.2.2.1 Topography and Bathymetry Topography influences the speed and direction of wind-wave and surge propagation, and frictional dissipation. Topography can also regionally amplify or attenuate storm surge. For these reasons, the topography in Coastal Texas was mapped in the STP 3 & 4 COLA ADCIRC hydrodynamic model using the most accurate and current topographic survey data (Reference 2.4-20), as obtained from a variety of data sources. The most recent and most trusted topographic values came from LIDAR data sets from the Texas Water Development Board (TWDB), Harris County Flood Control District (HCFCD), FEMA, and Louisiana State University (LSU)/Louisiana Oil Spill Contingency Office Atlas. The nodal elevations in the STP 3 & 4 COLA ADCIRC mesh were carefully reviewed at the interface between data sources and adjusted to smooth out any discontinuities. All data sets were converted to North American Datum of 1983 (NAD83) and elevations adjusted to NAVD 88.

Topographic data for the majority of the terrain in Texas were obtained from the TWDB (2007),

HCFCD (2002), FEMA (2006), and LSU LIDAR. These data were available in Digital Elevation Model (DEM) form on a 10-meter by 10-meter basis, and some later became available on a 1 meter by 1-meter basis. Small-scale hydro-dynamically relevant features, like levees, river banks, and road beds, are represented in the data. The STP 3 & 4 COLA ADCIRC mesh was primarily built using the 10-meter LIDAR because its nominal resolution best matched the intended resolution of most areas of the mesh. Some refinements of highly resolved, hydraulically relevant features were later made using the 1-meter LIDAR DEM. Figure 2.4-3 shows the topographic data sources near the STP site.

Aerial photography was not utilized as a source of bathymetric or topographic definition when creating the TX2008 Grid. However, once topography and bathymetry were defined using the most accurate data available, the horizontal alignment of major features like roadways, shorelines, and river banks was checked against aerial photographs using both proprietary and public source satellite images.

Accurate model bathymetry is crucial to accurately represent the flow physics of a region.

Bathymetry controls physical processes including long wave and short wind-wave propagation speed and direction, structure, and dissipation. Bathymetry in the regions in the western North Atlantic, Gulf of Mexico, and Caribbean Sea that are included in the models was drawn from a number of sources including the raw bathymetric sounding database from the National Ocean Service (NOS), the Digital Nautical Charts (DNC) bathymetric database, and ETOPO5 data from NOAA. The NOS raw sounding database provides the most comprehensive coverage over U.S. continental shelf waters, which include more that 13 million sounding values and is the basis of NOS/NOAA bathymetric charts.

Although the surveys are not as comprehensive as the NOS raw soundings, DNC values are available within the Gulf of Mexico and much of the western North Atlantic and Caribbean Sea, while ETOPO5 coverage is worldwide. Data accuracy and preferences are in the order NOS, DNC, and finally ETOPO5. Bathymetry for inland waterways in Coastal Texas is provided by regional bathymetric surveys and dredging surveys from the USACE Southwest Galveston District (SWG), NOAA, TWDB, or nautical charts. Figure 2.4-4 shows the source for STP COLA ADCIRC mesh bathymetric data for Matagorda Bay and the surrounding water bodies.

Storm Surge 2.4-6

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project 2.4.2.2.2.2 Model Grid Version 13 of the Texas topographic grid, herein referred to as TX2008 model, is an extension of the earlier EC2001 U.S. East Coast and Gulf of Mexico tide model and the TX04 Coastal Texas storm surge model. These models all incorporate the western North Atlantic Ocean, the Gulf of Mexico, and the Caribbean Sea to allow for full dynamic coupling between oceans, continental shelves, and the coastal floodplain without requiring definition of these complicated couplings in the boundary conditions. The TX2008 model extends the geographic coverage of these earlier models to include all the floodplains of Coastal Texas. In addition, improved feature definitions, surface roughness definition, wave radiation stress definition, and grid resolution were all incorporated into the TX2008 model.

The TX2008 model domain has an eastern open ocean boundary that lies along the 60-degree west meridian, extending south from the vicinity of Glace Bay in Nova Scotia, Canada, to the vicinity of Coracora Island in eastern Venezuela. This domain has a superior open ocean boundary that is primarily located in the deep ocean and lies outside of any resonant basin. This boundary has nominal geometric complexity. Tidal response is dominated by the astronomical constituents. Nonlinear energy is limited due to the depth. The boundary is not located near tidal amphidromes. Hurricane storm surge response along this boundary is essentially an inverted barometric pressure effect directly correlated to the atmospheric pressure deficit in the meteorological forcing. This boundary allows the model to accurately capture basin-to-basin and shelf-to-basin physics.

Open ocean bathymetric depths were first interpolated from a 5-degree by 5-degree regular grid based on the ETOPO5 values. The DNC bathymetric values were then applied over much of the Atlantic, Gulf of Mexico, and Caribbean. Bathymetric values were subsequently applied using the NOAA depth-sounding database. Thus, bathymetric values were applied with a priority/availability system with preference being given to the NOAA sounding database, then the DNC database, and then the ETOPO5 database. This preference is related to the accuracy of each database.

Much of the domain is bordered by a land boundary made up of the eastern coastlines of North, Central, and South America. The highly detailed/resolved region extends along the coast from Brownsville to Port Arthur, Texas. The coastal regions adjacent to Texas, northern Mexico, and western Louisiana were also included at high resolution in order to allow storm surge to realistically attenuate and laterally spread into the adjacent regions. In the Texas model, the domain includes a large overland region that is at risk for storm surge induced flooding.

Details of the domain with bathymetry and topography as well as levees and raised roadways across Coastal Texas in the vicinity of the STP site can be seen on Figures 2.4-5 and 2.4-6. The inland extent of the Texas model follows high topography or major hydraulic controls. The land boundary runs along the 30- to 75-foot land contour. The boundary was positioned such that lower-lying valleys and the adjacent highlands were included. It is critical that boundary location and boundary condition specification do not hinder physically realistic model response.

Critical hydraulic features and controls are included in the TX2008 grid that both enhance and attenuate storm surge. Rivers and channels can be conduits for storm surge propagation far inland. Topographical features such as levee systems stop flow and can focus storm surge energy into local areas, resulting in the amplification of storm surge. Floodplains and wetlands cause attenuation of flood wave propagation. Many interconnected features are in Texas, Storm Surge 2.4-7

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushinma Response Project including deep naturally scoured channels, wetlands, and an extensive and intricate system of river banks, levees, and raised roadways. Rivers, such as the Brazos River, Nueces River and Rio Grande, and numerous major dredged navigation canals including the Gulf Intracoastal Waterway (GIWW) are incorporated into the grid. All significant levee systems, elevated roads, and railways have been specifically included in the domain as barrier boundaries. These raised features are represented as a continuous row of elevated nodes, as internal barrier boundaries, or as external barrier boundaries when they are at the edge of the domain.

The mesh design provides localized refinement of the Texan coastal floodplains and of the important hydraulic features. The level of detail in Coastal Texas is very high, with nodal spacing reaching as low as 100 feet in the most highly refined areas. Unstructured meshes can resolve the critical features and the associated local flow processes with orders of magnitude of fewer computational nodes than a structured grid because the latter is limited in its ability to provide resolution on a localized basis, and fine resolution generally extends far outside the necessary area. The TX2008 mesh is refined locally to resolve features such as inlets, rivers, navigation channels, levee systems, and local topography/bathymetry. The TX2008 grid is shown in Figure 2.4-7.

The TX2008 computational mesh contains more than 2.8 million nodes and nodal spacing varies significantly throughout the mesh. Grid resolution varies from approximately 12 to 15 miles in the deep Atlantic Ocean to about 100 feet in Texas. The high grid resolution required for the study region leads to a final grid with more than 90 percent of the computational nodes placed within or upon the shelf adjacent to Texas, enabling sufficient resolution while minimizing the cost of including such an extensive domain. Use of a large-scale domain therefore only adds 10 percent to the computational cost of the simulations. The result, however, is the application of highly accurate boundary conditions and full dynamic coupling between all scales from basins to inlets.

2.4.2.2.2.3 Sub-Grid Scale Vertical Features Features such as barrier islands and river banks are generally well resolved in grids with resolutions down to about 100 feet. However, features like levees, floodwalls, railroads, and raised highways will not be sufficiently well resolved with 100-foot grid resolution. Frequently, these small-scale features can be significant horizontal obstructions to flow causing water to rise or be diverted elsewhere. These obstructions must therefore be carefully incorporated into the model as sub-grid scale weirs or lines of nodes specified as feature crown elevations. Their horizontal and vertical position must be well defined (see Figures 2.4-5 and 2.4-6). Sub-grid scale weirs were included with sub- and super-critical weir coefficients for features that were notably higher (i.e., 10 feet) than prevailing ground.

Federal, state, and local roads, and railroads and other continuous raised features, were positioned horizontally using LIDAR data or satellite images. Vertical positions were typically defined from the Texas 10-meter by 10-meter LIDAR data set. However, the elevations were also confirmed or adjusted with 1-meter by 1-meter LIDAR where available. The crown height was obtained automatically by searching a defined region around the raised feature's point of interest. Features were only included if the crown height was more than three feet above the adjacent topography and the feature was long enough to substantially impede flow.

Storm Surge 2.4-8

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 FukushinraResponse Project 2.4.2.2.3 Sensitivity of PMH Parameters on Storm Surge Elevation Combinations of PMH parameters listed in Table 2.4-1, PMH track direction and landfall location are used to specify the PMH scenarios that may occur at the STP site. Three individual values were selected for the PMH size, approach angle, forward speed and landfall location, resulting in a total of 81 PMH scenarios. The radius to maximum winds was set to 6, 12.9, and 20.8 miles (approximately 5, 11, and 18 nautical miles); the approach angle to 97.50, 143.80, and 190' clockwise from the north; and the forward speeds to 6.9, 14.4, and 21.8 miles per hour (approximately 6, 12.5, and 19 knots). Three landfall points were selected, with the first landfall point located at a distance equal to the radius of maximum winds, west of the mouth of the Colorado River Navigation Channel at the barrier islands. The second point was centered on the mouth of the Colorado River Navigation Channel at the barrier islands. The third point was located a distance equal to the radius of the maximum winds east of the mouth of the Colorado River Navigation Channel, at the barrier islands. All storm tracks were straight.

The results of storm surge simulations using STP 3 & 4 COLA ADCIRC model indicated that the maximum water surface elevation near the STP site would be produced by a large (in terms of radius to maximum winds), fast-moving (in terms of forward speed) storm that would produce prevailing winds blowing from the east or southeast toward the STP site. Because hurricanes rotate counter clockwise in the northern hemisphere, the highest surges are expected on the east side of the hurricane eye due to the fastest onshore wind being toward the right of the eye.

Storms with larger forward speeds generate faster responses in surge, leaving less time for the surge to dissipate over and around the surrounding terrain. Considering these factors, the site would be most vulnerable to flooding when the eye of the hurricane passes quickly to the west of the site on the leading edge of the storm. Based on the above outcomes and observations, it was concluded that the PMSS at the site would result from the PMH with the central pressure deficit and the upper bound radius of maximum winds and forward speed, as provided in Table 2.4-1.

Seven additional model sensitivity runs were performed for hurricane scenarios using the upper bound PMH parameters indicated above. The PMH parameters selected for the second set of STP 3 &4 COLA ADCIRC runs use a radius to maximum winds of 24 miles (21 nautical miles);

an approach direction of 1350 clockwise from the north (i.e. a northwesterly direction); a forward speed of 23 mph (20 knots); a central pressure of 26.19 in Hg; and a peripheral pressure of 30.12 in Hg. The only variables were the distance of the storm track from the site and the track direction. Table 2.4-3 lists the parameters for each of the seven simulated hurricanes. Figure 2.4-8 shows, as an example, the storm track and related STP 3 & 4 COLA ADCIRC inputs used for Scenario 2. As shown in the figure, steady state PMH conditions are considered prior to landfall and reduced hurricane intensities after landfall.

A sensitivity analysis was performed for the choice of the wind model to be used in STP 3 & 4 COLA ADCIRC. ADCIRC typically uses the Holland wind profile as the basis for calculating surface wind speeds as a function of distance from the storm center. However, for the same PMH attributes, wind profiles based on NWS 48 (Reference 2.4-21) generate a greater wind speed than the Holland wind profile. The ADCIRC code was therefore modified with the NWS 48 wind profile equation to generate the hurricane wind field.

Results of the second set of sensitivity simulations are summarized in Table 2.4-4. Scenario 2 could be identified as the PMH condition that would generate the PMSS elevation at the site.

Storm Surge 2.4-9

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 Fukushima Response Project 2.4.2.2.4 Maximum Storm Surge and Wind Speed for the Selected PMH Based on the results presented in Table 2.4-4, the PMSS elevation at the STP site is 29.3 ft NGVD 29 (or MSL). This PMSS will occur as the result of a hurricane traveling towards northwest direction (i.e., an approach direction of 1350 clockwise from the north) passing within 24 miles of the STP site. During its life up to the point of landfall, the storm will have a constant forward speed of 23 miles per hour, a central barometric pressure of 887 millibar, and a maximum sustained wind speed (1-mmin average) of 160 knots (184 miles per hour). Upon landfall the storm will continue in the same direction, but would begin to decay gradually as it moves inland. The simulated storm surge elevation for the selected set of PMH parameters for Scenario 2 is shown on Figure 2.4-9.

2.4.2.2.5 Uncertainties in Predicted PMSS Elevation Resio et al. (Reference 2.4-22) identified at least four sources of uncertainties in estimating very low probability hurricanes, like the PMH. These are:

" Uncertainty in defining hurricane intensity;

Uncertainty in storm surge model prediction;

" Potential climate variability over the projected design life; and

" Coincidental tide levels.

The last two items are accounted for in the STP 3 & 4 COLA ADCIRC model by incorporating a long term sea level rise of 1.43 ft for a 100-year period and the 10 percent exceedance high spring tide in the antecedent sea level, per Reference 2.4-19. The PMH intensity, which is the same as the pressure deficit, is adopted from NWS-23 (Reference 2.4-6). As described in Subsection 2.4.1, NWS-23 includes the past active hurricane periods since 1851 in developing the PMH parameters. Therefore, it is reasonable to assume that the PMH parameters from NWS-23 would remain reasonably conservative in the consideration of future climate variability.

Regarding uncertainties in model prediction, the ADCIRC model as applied to the STP storm surge analysis underwent an extensive flood level evaluation process to validate model performance over a range of conditions to ensure that the flow physics of the system were accurately characterized. The set of validation storms specific to the Texas coastal areas included Hurricanes Carla (1961), Celia (1970), Allen (1980), Alicia (1983), Bret (1999), Rita (2005), and Ike (2008). Hurricanes Rita and Ike were particularly useful storms for validation because of the large magnitude of surge they produced, and the accurate measurements of wind, atmospheric pressure, waves, and surge levels that exist for these two storms. Also, the TX2008 grid used in this application uses the most updated and detailed set of topographic information that resulted in a highly resolved grid region extending well inland near the STP site.

Resio et al. (Reference 2.4-22) provided a quantitative evaluation of uncertainties related to storm surge prediction based on the comparison of ADCIRC model results against observed storm surge elevations for Hurricane Katrina in the Louisiana coast. The comparison (as shown in Figure 4-17 in Reference 2.4-22) shows that the model generally over predicted the storm surge near high surge levels. This is also evident as the best fit between simulated and observed surge elevations shows a bias towards simulated storm surge for higher water levels.

A similar trend in model prediction can also be observed for SLOSH (Sea, Lake, and Overland Surges from Hurricanes) in the NOAA Technical Report NWS-48 (Reference 2.4-21).

Storm Surge 2.4-10

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushitna Response Project Given that the ADCIRC model used for the STP site was validated against historical hurricanes with landfall on the Texas coast and that comparison of observed storm surge levels with simulated results showed generally higher simulated surge level near the highest recorded water levels, it can be concluded that no further adjustment to the computed PMSS elevation to account for modeling uncertainties is warranted. Consequently, the STP 3 & 4 COLA ADCIRC model simulated surge elevation of 29.3 ft MSL is considered as the appropriate PMSS still water level for the STP site.

2.4.3 Wave Actions The effects of the PMH wind field and coincidental wind-wave actions on the STP 1 & 2 safety-related SSCs are evaluated to estimate the maximum flood level during a PMH event.

2.4.3.1 Hurricane Maximum Wind Speed The maximum 1-minute average, 10-m (33-foot)-high sustained wind speed at the STP site used in the STP 3 & 4 COLA ADCIRC model is 184 mph, which is used to calculate the coincidental wind-wave runup.

2.4.3.2 Wave Height, Period and Run-up on Safety Related SSCs The wind setup due to the PMH wind field and wave setup are included in the surge elevation obtained in the STP 3 & 4 COLA ADCIRC model results. However, wind-waves driven by the PMH wind and coincidental with the PMSS elevation would produce wave run-up on STP 1 & 2 safety-related SSCs.

In the power block, the grade level is at approximately 28.0 ft MSL. At the reevaluated PMSS still water level of 29.3 ft MSL, the maximum inundation flood depth would be shallow at about 1.3 ft during a PMH event. The minimum CDB flood elevation for the safety-related SSCs in the power block area at 44.5 ft MSL (Section 1.2) offers a substantial margin of over 15 ft over the PMSS still water level. In addition, the power block SSCs are surrounded by concrete vehicular barriers (of about 3.5 ft high sitting at grade), roads and other building structures that provide a shelter from direct wave actions and limit wave growth during the passage of a PMH. It can therefore be concluded that the safety-related SSCs in the power block would not be adversely affected by the wave actions coincidental with the PMSS still water level of 29.3 ft MSL.

During a PMH event, the maximum flood level on the ECWIS, which is a safety-related structure located on the ECP embankment, would be due to the wind setup and wave runup as a result of the hurricane wind acting on the stored water inside the ECP. The maximum runup level on the ECWIS is predicted to be 40.1 ft MSL as described in Section 2.5.

Figure 2.4-10 shows the ECP and elevation contours around the ECP. Because the area south and southeast of the ECP would be inundated during the PMSS event, the fetch in this direction is assumed conservatively to be unlimited. With the PMH forward speed after landfall at about 20 mph (as shown in Figure 2.4-8) and the site located approximately 16 miles from the Gulf of Mexico shoreline, the maximum wind would be sustained for less than one hour over the fetch.

Considering a duration limited wave for a maximum period of one hour, the deep water significant wave height and peak period can be estimated as 13.9 ft and 4.9 seconds, respectively, following the methodology described in the Coastal Engineering Manual (CEM)

Storm Surge 2.4-11

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP I & 2 Fitkushima Response Project (Reference 2.4-23). The wave height would reduce due to the breaking process before reaching the ECP embankment because of the limited water depth.

The grade elevation is about 27 ft MSL at the foot of the ECP embankment, and varies between 25 and 27 ft MSL over a wide area further southeast, as shown in Figure 2.4-10. The water depth of 2.3 ft at the exterior toe of the ECP embankment would sustain a significant and maximum breaking wave height of approximately 1.4 and 1.9 ft, respectively.

Wave runup on the ECP embankment, which has a side slope of 1V:3H and a crest elevation of 34 ft MSL, is also estimated based on the methodology provided in the CEM (Reference 2.4-23). The significant, 2 percent, and maximum wave runup on the ECP embankment are estimated at 3.2 ft, 4.4 ft and 5.9 ft, respectively. The storm surge level including the significant, 2 percent and maximum wave runup on the ECP embankment would be 32.5 ft MSL, 33.7 ft MSL and 35.2 ft MSL, respectively. While the water level associated with the significant and 2 percent wave runup would remain below the ECP embankment crest, the water level associated with the maximum wave runup would exceed the ECP embankment crest elevation. However, because of the 4.7 ft freeboard, defined as the height of the ECP embankment crest above the PMSS still water level, the overtopping rate on the ECP embankment for the duration of the PMSS would be small, on the order of 2.81x104 ft3/s/ft, and would have negligible effect on the water level inside the ECP. For instance, assuming conservatively that this maximum overtopping rate would last for an hour of duration and over an embankment length of 2000 ft as shown in Figure 2.4-10, the increase in water level would be less than 0.001 ft based on a surface area of 46.5 acres at the ECP normal operating water level of El. 25.5 ft MSL (Reference 2.4-1, Subsection 9.2.5.2) The maximum runup level at the ECWIS as a result of PMH wind action, in combination with a 100-year 4-day precipitation, is predicted to be 40.1 ft MSL as part of seiche flooding evaluation described in Section 2.5.

The cooling water discharge structure is located on the west ECP embankment as the ECWlS but on the north side of the dividing dike. It has a lower profile than the ECWIS, with the top elevation at 36 ft MSL. The 40.1 ft MSL maximum wave runup level at the ECWIS will provide a bounding flood level for the discharge structure as well during a PMH event.

PMSS flooding impact was evaluated for the STP 1 & 2 site as described in the UFSAR (Reference 2.4-1). The PMSS still water level was estimated at 26.78 ft MSL, which is lower than the reevaluated level of 29.3 ft MSL. With the power block grade elevation of 28 ft MSL, the UFSAR concluded that the PMSS of 26.78 ft MSL would not inundate the STP site and surrounding areas. Consequently, the coincidental wave runup level is not defined in the UFSAR for the safety-related SSCs. In comparison, the maximum reevaluated flood elevation including wave runup at the ECP embankment exterior is estimated as 35.2 ft MSL, as demonstrated in this section, which could also conservatively be adopted for the power block safety-related SSCs. The maximum flood elevation at the ECWIS and the discharge structure is reevaluated as 40.1 ft MSL.

2.4.4 Protective Structures The ECWIS withdraws water from the ECP, which is confined by an embankment system. As described in Subsection 1.1.2, the embankment system includes safety-related and nonsafety-related portions. The safety-related portion of the ECP embankment is located along the west and south of the ECP. The ECP embankment has a crest elevation of 34 ft MSL and includes a 30-ft wide interior berm at 26 ft MSL. Both sides of the safety-related ECP embankment are Storm Surge 2.4-12

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STPI & 2 Fukushima Response Project protected against erosion. As described in Subsection 2.5.6.4.2.4 of the STP Units 1 & 2 UFSAR (Reference 2.4-1), slope protection for the south and west (safety-related) Category I embankment sections is provided by 6 in. of reinforced concrete. The same type of concrete protection is provided for the interior berm of the Category I sections, the dike berms, and local areas of the 5:1 slope of the pond excavation adjacent to intake and discharge structures. Slope protection for the reminder of the interior berm of the embankment, and the 5:1 slope of the pond excavation, is provided by a 1.25-ft layer of soil-cement.

Because the PMH that passes by the west of the site would generate the PMSS at the site, the safety-related portion of the south ECP embankment would be most exposed to direct wave actions. The current embankment design with concrete and soil cement linings would provide sufficient protection against wave and current erosion. In addition, overtopping at the embankment during the PMSS will not result in any adverse safety impact as the overtopping flow is minimal and occurs intermittently only at the maximum wave height in the wave train as demonstrated above.

Debris may accumulate during the PMSS near the ECP. However, due to the presence of the ECP embankment, only floating debris could enter the ECP during maximum runup condition, which produces a minimal overtopping flow into the ECP. It is unlikely that such floating debris would affect the ECP embankment or the ECWIS. In addition, the safety related ECW pumps housed in the ECWIS structure are protected from influx or debris by track bars and traveling water screens. The cooling water discharge structure is a passive structure and its operation would not be affected by flood inundation, wave runup or overtopping. The discharge outlets are protected by security bars that will prevent carrying over of large debris of concern with the waves into the discharge pipes creating blockage problems.

The safety-related SSCs in the STP site are designed to withstand the hydrostatic and hydrodynamic forces from a postulated MCR embankment breach event, which is the controlling flood event for the STP 1 & 2 safety related facilities and produces higher flood levels than the PMSS condition. Consequently, hydrostatic and hydrodynamic forces from the PMSS and coincidental wave actions on the safety SSCs would be bounded by the MCR embankment breach event.

2.4.5 References 2.4-1 STPEGS (Units 1 & 2) Updated Final Safety Analysis Report (UFSAR), Subsections 2.4.4 and 2.4.5, Rev. 15; Subsections 2.5.6 and 9.2.5, Rev. 16.

2.4-2 South Texas Project Units 3 &4 Combined Licensing Application (COLA), Final Safety Analysis Report (FSAR), Rev. 7, Nuclear Innovation North America LLC, February 1, 2012.

2.4-3 "Guidance for Performing a Tsunami, Surge, or Seiche Hazard Assessment", U.S.

Nuclear Regulatory Commission, Interim Staff Guidance, Rev. 0, JLD-ISG-2012-06, January 2013.

2.4-4 "Determining Design Basis Flooding at Power Reactor Sites", ANSI/ANS-2.8-1992, American Nuclear Society, July 1992. (Historical Technical Reference).

Storm Surge 2.4-13

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STPI & 2 Fukushima Response Project 2.4-5 "Design-Basis Flood Estimation for Site Characterization at Nuclear Power Plants in the United States of America", NUREG/CR-7046, U.S. Nuclear Regulatory Commission, November 2011.

2.4-6 "Meteorological Criteria for Standard Project Hurricane and Probable Maximum Hurricane Windfields, Gulf and East Coast of the United States," National and Atmospheric Administration (NOAA) Technical Report NWS 23, Schwerdt, R. W.,

NOAA, 1979.

2.4-7 National Oceanic and Atmospheric Administration (NOAA), FAQ/State of the Science: Atlantic Hurricane & Climate, U.S. Department of Commerce, December 2006.

2.4-8 Blake, E.S., et al., The Deadliest, Costliest, and Most Intense United States Tropical Cyclones from 1851 to 2010 (and Other Frequently Requested Hurricane Facts),

Technical Memorandum NWS TPC-6, National Weather Service, National Hurricane Center, National Oceanic and Atmospheric Administration (NOAA), August 2010.

2.4-9 "SLOSH: Sea, Lake, and Overland Surges from Hurricanes," National and Atmospheric Administration (NOAA) Technical Report NWS 48, Jelesnianski, C. P.,

Chen, J. and W. A. Shaffer, NOAA 1992.

2.4-10 "Storm Surge on the Open Coast: Fundamentals and Simplified Prediction,"

Technical Memorandum No. 35, Bodine, B.R., U.S. Army Corps of Engineers Coastal Engineering Research Center, 1971.

2.4-11 "Storm tracks for Atlantic Basin," National Oceanic and Atmospheric Administration.

Available at http://maps.csc.noaa.gov/hurricanes/download.html, accessed May 4 2007.

2.4-12 "Hurricane Surge Frequency Estimated for the Gulf Coast of Texas," Technical Memorandum No. 26, Bodine, B. R., U.S. Army Corps of Engineers, Coastal Engineering Research Center, 1969.

2.4-13 "Verification Study of a Bathystrophic Storm Surge Model," Technical Memorandum No. 50, Pararas-Carayannis, George, U.S. Army, Corps of Engineers - Coastal Engineering Research Center, May 1975.

2.4-14 "Design and Implementation of a Real-Time Storm Surge and Flood Forecasting Capability for the State of North Carolina," Mattocks et al, Carolina Environmental Program, University of North Carolina, November 30, 2006 2.4-15 "A Basin- to Channel-Scale Unstructured Grid Hurricane Storm Surge Model Applied to Southern Louisiana," Westerlink et al, American Meteorological Society, March 2008.

2.4-16 "A High-Resolution Coupled Riverine Flow, Tide, Wind, Wind Wave, and Storm Surge Model for Southern Louisiana and Mississippi. Part I: Model Development and Validation," Bunya et al, American Meteorological Society, February 2010.

Storm Surge 2.4-14

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project 2.4-17 "A High-Resolution Coupled Riverine Flow, Tide, Wind, Wind Wave, and Storm Surge Model for Southern Louisiana and Mississippi. Part I1: Synoptic Description and Analysis of Hurricanes Katrina and Rita," Dietrich et al, American Meteorological Society, February 2010.

2.4-18 "Modeling Hurricane Waves and Storm Surge using Integrally-Coupled, Scalable Computations," J.C. Dietrich et al., Coastal Engineering, July 9, 2010.

2.4-19 "Design Basis Floods for Nuclear Power Plants," Regulatory Guide 1.59, Revision 2, U.S. Nuclear Regulatory Commission, 1977.

2.4-20 "Flood Insurance Study: Coastal Counties, Texas, Intermediate Submission 1:

Scoping and Data Review," 7 January 2010, USACE New Orleans District and FEMA Region 6.

2.4-21 Jelesnianski, C.P., et al., SLOSH: Sea, Lake, and Overland Surges from Hurricanes, Technical Report NWS 48, National Oceanic and Atmospheric Administration (NOAA), April 1992.

2.4-22 Resio, D.T., et al., The Estimation of Very-Low Probability Hurricane Storm Surges for Design and Licensing of Nuclear Power Plants in Coastal Areas, NUREG/CR-7134, U.S. Nuclear Regulatory Commission, October 2012.

2.4-23 US Army Corps of Engineers, Coastal Engineering Manual, EM 1110-2-1100 Part II:

Chapter 1, Chapter 2, 1 August 2008; Chapter 4, 31 July 2003; Part VI: Chapter 5, 28 September 2011.

Storm Surge 2.4-15

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STPI & 2 Fukushima Response Project Table 2.4-1 Probable Maximum Hurricane Characteristics Peripheral Pressure (pn) 30.12 in. Hg.

Central Pressure (Po) 26.19 in. Hg.

Radius of Maximum Winds (R) 5 to 21 nautical miles Forward Speed (T) 6 to 20 knots Notes: The parameters given in the table are: R radius of maximum winds:

T translation speed; Vxsmaximum stationary wind speed Storm Surge 2.4-16

Enclosure NOC-AE-13002975 FloodingHazardReevaluation Report STP1 & 2 FukushirnaResponse Project Table 2.4-2 Major Historic Hurricanes Impacting the Texas Coast 1900 to 2005 Date of Landfall Namne 24 Sep 2005 Hurricane Rita 24 Sep 2004 Hurricane Ivan 15 Jul 2003 Hurricane Claudette 23 Aug 1999 Hurricane Bret I Aug 1989 Hurricane Chantal 16 Oct 1989 Hurricane Jerry 18 Sep 1988 Hurricane Gilbert 26 Jun 1986 Hurricane Bonnie 18 Aug 1983 Hurricane Alicia 10 Aug 1980 Hurricane Allen 8 Sep 1974 Hurricane Carmen 10 Sep 1971 Hurricane Fern 3Ocg 1970 Hurricane Celia 20 Sep 1967 Hurricane Beulah 17 Sep 1963 Hurricane Cindy 3 uSep 1961 Hurricane Carla 25 Jun 1959 Hurricane Debra 4 Oct 1949 1949 Hurricane 27 Aug 1945 1945 Hurricane 27 Jul 1943 1943 Hurricane 30 Aug 1942 1942 Hurricane 23 Sep 1941 1941 Hurricane 7 Aug 1940 1940 Hurricane 1 Sep 1933 1933 Hurricane 4 Aug 1933 1933 Hurricane 13 Aug 1932 1932 Hurricane 23 Jun 1929 1929 Hurricane 21 Jun 1921 192 1 Hurricane 14 Sep 1919 1919 Hurricane 18 Aug 1916 1916 Hurricane 16 Aug 1915 1915 Hurricane 2 1 Jul 1909 1909 Hurricane 8 Sep 1900 Galveston Hurricane Source: References 2.4-11 and 2.4-12 (Hurricanes Dolly and Ike made landfall in July and September, 2008. respectively, are not included in this table)

Storm Surge 2.4-17

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STPI & 2 Fukushima Response Project Table 2.4-3 Probable Maximum Hurricane Parameters and ADCIRC Model Scenarios Scnro12 3 4 5 6 1 Central 887 Mb 887 Mb Same Same Same Pressure (26.19 Hg) (26.19 Hg)

Peripheral 1020 Mb 1020 Mb Same Same Same Pressure (30.12 Hg) (30.12 Hg)

Pressure 133 Mb 133 Mb Same Same Same Differential (3.93 Hg) (3.93 Hg)

Radius to 24 miles 24 miles Maximum 24 mi Same Same 24 mi Same Same Same Winds (21 nm) (21 nm)

Forward 23 mph Same Same 23 mph Same Same Same Speed (20 knots) (20 knots)

Maximum 184 mph 184 mph Sustained Same Same Same Same Same Wind Shortest 12 miles 24 miles 36 miles 24 miles Distance 12 mi 24 mi 3 me 24 ml Same Same Same from site (10.4 nm) (20.9 nm) (31.3 nm) (21 nm)

Track NS e a Direction NW Same Same Storm Surge 2.4-18

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 Fukushima Response Project Table 2.4-4 Summary of ADCIRC Results for the Selected Scenarios 4 5 a7 Shortest 24 (21miles Distance 12 miles 24 miles 36 miles mil Same Same (21 nm) Same from site (10.4 nm) (20.9 nm) (31.3 nm)

Track NW Same Same Direction Maximum 25.0 29.0 26.0 20.0 26.5 29.3 28.5 Surge Height, ft (8.08 m) (8.93 m) (8.67 m) (7.62 m) (8.84 m) (7.92 m) (6.10 m)

Storm Surge 2.4-19

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 0 OID XDX0~ AM. Bagwnetfc cont"u. m - -- Humcaneftrack Source: Reference 2.4-11 Figure 2.4-1 Historic Hurricane Tracks of Major (i.e., Category 1 and Larger) Unnamned Hurricanes Impacting the Texas Coast Between 1852 and 1950 Storm Surge 2.4-20 I

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project iN

ýW SW A *D2MO Bstryetnc coftour~m - Humcane track Source: Reference 2.4-11 Figure 2.4-2 Historic Hurricane Tracks of Major (i.e., Category 1 and Larger)

Unnamned Hurricanes Impacting the Texas Coast from 1950 to 2006 Storm Surge 2.4-21 I

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STPI & 2 Fukushima Response Project

/

i L"WW 0

0 TOPOW

ý fC- kJ ýW( .X" kIr ! *~.wný. '(0 gn,!

Figure 2.4-3 Topographic Data Sources for the TX2008 Grid near the STP Site Storm Surge 2Z4-22

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 FukushimaResponse Project

- 'C RAS Ma" Dom MOAA-%OS ""WW~

46 OD" S*Utg

-TfO4IIuI AOcCW Figure 2.4-4 Bathymetric Data Sources for the TX2008 Grid near the STP Site Storm Surge 2.4-23

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project 281 28; 28E 3

28!

284-28 3 282-

-W I -VDU Longftft (do-Figure 2.4-5 Topographic Features of the TX2008 Grid Near the STP Site Figure 2.4-6 Landward Extent of TX2008 Grid near the STP Site Storm Surge 2. 4-24

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluationReport STP I & 2 Fukushima Response Project Figure 2.4-7 The TX2008 Grid Storm Surge 2.4-25

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project tam" Uala C ooleasb -sele sosws isFa owie b "Wu of C*UY fgf NORTHWEST 1 3 27.72 95.52 5 87 24 60 69 23 20 2 6 X69 94.73 5 887 24 60 99 23 20 PMH Storm Features 3 9 26.27 93.96 5 87 24 60 69 23 20 4 12 2".54 93.19 5 887 24 60 69 23 20 Ce..94 Plessixe: W Mb 26.19 In. Vg 5 15 24.81 92,43 887 24 60 O9 23 20 Peripheral Pr~esure I0M Mb (30.12in. V 6 18 2409 91.67 3 88 24 40 69 23 20 Radkis to MaxinveisnWNK Z r1um, 124miles) 7 21 23.37 90.91 4 944 20 60 69 23 20 Fouwwd Speedt 20ht (23 m~ph) 8 24 22M 90.12 3 964 16 60 69 23 20 Mask.... Sustained 1101111 160knob (194 ph) 9 27 21.89 89.34 2 979 13 Shortest 011stmtcm fbrogsite: 2.9 am. 1Nen006M Figure 2.4-8 PMH Track and STP COLA ADCIRC model Parameters for Scenario 2 Storm Surge 2.4-26

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project 29 ft 34 32 30 28 26 24 22 20 18 16 14 12 10 8

6 4

2 28" 0

-96' -95" Figure 2.4-9 STP COLA ADCIRC Model Results for the Maximum Water Level for Scenario 2 Storm Surge 2.4-27

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Figure 2.4-10 Elevation Contours (ft MSL) Surrounding the ECP Storm Surge 2.4-28

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 FukushimaResponse Project 2.5 Seiche A seiche is a standing wave in an enclosed or partially enclosed body of water which can be caused by external forcing mechanisms such as wave, tidal, atmospheric and seismic forcing.

When the forces causing the wind/pressure changes stop, seiching oscillations on the water surface might appear. The frequency of oscillation is a function of geometry and bathymetry of the water body.

A site specific evaluation is conducted in accordance with the guidance in NUREG/CR-7046 (Reference 2.5-1) and ANSI/ANS 2.8-1992 (Reference 2.5-2) to examine the flooding potential at STP 1 & 2 as a result of seiche events from different initiators. In particular, the potential of seiching oscillations in the water bodies at or adjacent to the site triggered by the passage of a moving front or a hurricane is evaluated.

The two enclosed water bodies at or near STP I & 2 that would potentially subject to seiche motions are the Main Cooling Reservoir (MCR) with a nominal surface area of 7000 acres (Reference 2.5-3, Subsection 2.4S.1.1), and the Essential Cooling Pond with a surface area of 39.2 acres at El. 17 ft MSL and 46.5 acres at El. 25.5 ft MSL (Reference 2.5-4, Subsection 9.2.5.2). The other major water bodies in the vicinity of STP 1 & 2 are the Lower Colorado River, Matagorda Bay and the Gulf of Mexico. The seiche induced flooding potential in each of these water bodies are described in the following subsections.

2.5.1 Seiches in the Essential Cooling Pond To assess the potential flood risk due to seiching motions induced by wind and pressure systems passing over the ECP, the wind setup, surface wave characteristics and corresponding runup at the ECP and the Essential Cooling Water Intake Structure (ECWIS) are estimated for the critical combined event conditions specified in ANSI/ANS 2.8-1992 (Reference 2.5-2). The natural frequencies of the ECP are also evaluated for resonance effects.

In accordance with Reference 2.5-2, wind setup, wave runup and potential seiche flooding are evaluated specifically for the STP 1 & 2 safety-related ECP for two conditions: (1) a 2-year wind event with a water level based on the combination of the operating water level and the site Probable Maximum Precipitation (PMP) to assess the wind induced wave and setup and the freeboard above still water level of a safety-related reservoir; and (2) the Probable Maximum Hurricane (PMH) passing over the ECP containing its 100-year water level or maximum controlled level in water body, whichever is less. The latter condition is to evaluate the surge and seiche effects at shore locations on enclosed bodies of water.

Figure 2.5-1 shows a typical cross-section profile of the ECP. The ECP normal operating water level varies between 25.6 ft and 26.0 ft MSL (Reference 2.5-4, Subsection 9.2.5.2). Given the ECP bottom elevation of 17 ft MSL, the ECP water depth during normal operating condition varies between 8.6 ft and 9 ft.

The PMP depth at the site used in the evaluation is 60 in (5 ft) (Table 2.5-1 and Figures 2.5-3a and 2.5-3b), for a 96-hour storm duration, derived based on an extrapolation from rainfall depths of shorter durations, 6-hour to 72-hour. The shorter duration PMP depths are obtained from NOAA's National Weather Service (NWS) Hydrometeorological Reports (HMRs) No. 51 and 52 (References 2.5-5 and 2.5-6). The maximum water depth in the ECP during the 96-hour PMP event would therefore be about 14 ft.

Seiche 2.5-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 FukushiniaResponse Project The 100-year water level at the ECP is determined using the isopluvial map of 100-year 4-day precipitation developed in Technical Paper 49 of NWS (formerly US Weather Bureau)

(Reference 2.5-7). A 100-year rainfall depth of about 16 in (1.3 ft) is estimated for the STP site, resulting in a water depth of about 10.3 ft at the ECP.

Seiching period can be estimated using the formula below applicable to a closed rectangular basin, with a constant depth, a reasonable assumption for the ECP due to its relatively constant depth and regular geometry:

T21b where g is the gravitational acceleration; T, /b and h are natural period, length and depth of the basin, respectively (Reference 2.5-8, Section 11-5-6).

2.5.1.1 Wind Setup, Waves Characteristics and Wave Runup in the ECP The ECP is formed by a perimeter embankment with a crest elevation of 34 ft MSL. There is a dividing dike with a crest elevation of 38 ft MSL (Reference 2.5-4, Subsection 2.4.4) that runs between the ECWIS and the discharge structure, separating the ECP into 2 sections. The longest fetches within the ECP that could potentially induce the largest waves at each of the sections are identified and shown in Figure 2.5-2.

The wind setup and wave runup at the ECP are determined, based on the methodology recommended by the U.S. Army Corps of Engineers (USACE) Coastal Engineering Manual (CEM) (Reference 2.5-8) and the other supplementary references (References 2.5-9 and 2.5-10).

According to ANSI/ANS 2.8-1992 (Reference 2.5-2), the 2-year mean recurrence interval annual fastest-mile wind speed measured 30 ft above the ground at the STP site is 50 mph.

Based on the storm surge reevaluation described in Section 2.4, the PMSS as generated by the ADCIRC model, conducted in support of the STP 3 & 4 Combined License Application (COLA),

is estimated to be 29.3 ft MSL. This PMSS will occur as the result of a hurricane traveling in a northwesterly direction (i.e., an approach direction of 135' clockwise from the north) passing within 24 miles of the STP site. Up to the point of landfall, the storm will have a constant forward speed of 23 mph, a central barometric pressure of 887 Mb, and a maximum sustained wind speed of 160 knots (184 mph). Upon landfall the storm will continue in a northwesterly direction, but began to decay gradually as it moves inland. Therefore, the STP PMH overwater maximum sustained wind speed (10 m, 1-min averaged) of 184 mph is conservatively chosen for the combined PMH and 100-year water level for the ECP. These wind speeds are adjusted for duration, overwater, air-water temperature difference and fetch length, as applicable, before being used to determine the wave characteristics.

The local intense precipitation flooding analysis for the site indicates that the peak water level around the site varies between 31.6 ft MSL and 33 ft MSL (Section 2.1, Table 2.1-6). Local PMP flooding typically has short durations, and considering also the blockage of the wind by the buildings, barriers and land features surrounding the ECP and the steeper ground surface slope Seiche 2.5-2

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project toward the Lower Colorado River to the east of the site, wave overtopping the embankment crest into the ECP during the PMP flooding would be negligible.

For the combined 2-year wind and PMP event, the maximum wind setup of approximately 0.04 ft, resulting in a maximum still water elevation of approximately 31.0 ft MSL, is estimated for the ECP. For this event, the significant and maximum wave heights and peak wave peak period, at the toe of the ECP embankment, are estimated to be 1.1 ft, 2.0 ft and 1.5 sec, respectively.

Such wave induces a maximum wave runup of 2.5 ft at the ECP embankment. Accordingly, the ECP embankment maximum water level, including the wind setup and wave runup, is estimated to reach at about 33.5 ft MSL. At the ECWIS, the combined 2-year wind and PMP event would potentially cause a maximum wave runup of about 3 ft, resulting in a maximum water elevation of about 34.0 ft MSL. The wave runup at the ECWIS is estimated using Goda's formula, given in Reference 2.5-10, that utilizes the maximum wave height and direction to determine the maximum wave runup on a vertical wall. The incident wave angle is conservatively assumed to be perpendicular to both ECP embankment and the ECWIS.

Section 2.4, Storm surge for flood hazard reevaluation, concludes that during the PMH event, the ECP water level rise due to the wave overtopping rate into the ECP from outside is negligible during the PMSS at the site. Therefore, for the combined PMH and the ECP 100-year water level event, the maximum still water elevation of about 27.9 ft MSL is estimated based on the maximum normal operating level of 26 ft MSL, 100-year 4-day precipitation of about 1.3 ft and the maximum wind setup of about 0.6 ft. For this event, the significant and maximum wave heights and peak wave period, at the toe of the ECP embankment, are determined to be 1.3 ft, 1.7 ft and 2.3 sec, respectively. Such wave generates a maximum wave runup of 4.0 ft at the ECP embankment. As a result, the ECP embankment maximum water level, including the wind setup and wave runup, is estimated to reach at 31.9 ft MSL. The combined PMH and the ECP 100-year water level event would potentially generate a maximum wave runup of about 12.2 ft, resulting in a maximum water elevation of about 40.1 ft MSL at the ECWIS, based on the conservative assumption of direct wave attack angle. This maximum water elevation is lower than the current design basis flood elevation of 40.8 ft MSL at the ECWIS documented in the UFSAR.

2.5.1.2 Resonance Effects in ECP due to Atmospheric Forcing Using the equation given in Subsection 2.5.1, the natural period of the ECP with a depth ranging between 8.6 ft and 14 ft, corresponding to the minimum operating water level and the combined maximum operating water level and the PMP, is estimated to be in the range of 3.7 min to 4.7 min.

In order for seiching oscillations to resonate, the forcing period should be equal or very close to the seiching period of the ECP. Given that the wave periods, which are on the order of a few seconds as estimated in Subsection 2.5.1.1, are significantly smaller than the ECP natural periods, it is unlikely for those oscillations to be resonated by the short waves at the ECP.

Atmospheric forcing events such as cyclical wind and moving pressure systems typically have a time scale on the order of hours, appreciably longer than the ECP natural periods. Therefore, it is highly unlikely for any atmospheric forcing to be capable of amplifying seiche motions in the ECP to cause flooding concerns at the ECW safety-related facilities.

In addition, the energy of the oscillations induced by the wind setup dissipates due to frictional effects. As a result, the seiching oscillations initiated by the wind setup are gradually damped.

Seiche 2.5-3

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STPI & 2 Fukushinta Response Project Conservatively assuming a negligible energy dissipation rate for such waves to travel across the ECP, the maximum elevations that the wind setup can reach at various locations inside the ECP are estimated to be approximately 31.0 ft MSL during the combined 2-year wind and PMP event and about 27.9 ft MSL during the combined PMH and 100-year water level event.

2.5.2 Seismic Seiche Typical seismic wave periods are on the order of seconds making it unlikely for the ECP water body seiching to be amplified from seismic motions because of the significant difference between the forcing period and the ECP's natural period.

The only documented event of a seismic seiche on the Texas coast is coincided with the 1964 Alaska (Mw=9.2) earthquake located between the Aleutian Trench and the Aleutian Volcanic Arc (Reference 2.5-12). The event was recorded on a tide gage in Freeport, Texas (Reference 2.5-13). Reference 2.5-13 (p. 261) notes that "in several reports from eyewitnesses in the coastal regions of Louisiana and Texas, waves up to 6 feet (2 meters) in height were observed."

However, Reference 2.5-13 (p. 261) reports that the "maximum height of the recorded seiche at 0400 GMT is about seven inches (18 cm)," and that the "true wave height may have been several feet ([i.e.,] about a meter)." Additional analyses of tide gage records from the 1964 event report the maximum measured height of the low-frequency waves along the Texas coast from the Alaska earthquake ranged from 0.22 to 0.84 feet (Reference 2.5-14, p. 26). Reference 2.5-14 indicated that the horizontal acceleration associated with seismic surface waves from the Alaska shock appears to have varied markedly within North America. The amplitude of horizontal acceleration was especially large along the Gulf coast. Reference 2.5-14 (p. 27) further stated that "thick deposits of sediments of low rigidity along the Gulf coast, for example, are capable of amplifying the horizontal acceleration of surface waves to a considerable extent; this accounts for the concentration of seiches that occurred along the Gulf coast."

While the M,=9.5 magnitude 1960 earthquake in Chile might also have been expected to have caused seiches along the Texas coast, tide gages along the Gulf coast did not record any event. The Mw=7.8 New Madrid earthquake that occurred on February 7, 1812 (Reference 2.5-15), which is the largest earthquake recorded in the contiguous United States, produced significant seiches in the Mississippi River and in waterways along the Texas state boundary (Reference 2.5-16, p. 124). However, no records exist to indicate that the 1812 New Madrid earthquake directly affected the South Texas coast or the Lower Colorado River near STP 1 &

2.

Further descriptions on historical information and data on tsunami-generating earthquakes and runup events are provided in Section 2.6.

2.5.3 Seiche in Main Cooling Reservoir A quantitative analysis of the seiche motions and overtopping potentials in the MCR is not conducted specifically in this evaluation because MCR is not a safety related facility and the loss of cooling function as a result of seiche flooding would not affect the safety of the plant.

Further, in the unlikely event of overtopping at the MCR embankment as a result of seiche motions, the flooding impacts would have been bounded by those resulting from the failures of the MCR embankment, the design basis flooding mechanism of STP I & 2, as described in Subsection 2.3.2.

Seiche 2.5-4

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushimna Response Project In addition, a recent seiche analysis conducted in support of the Combined License Application for STP 3 & 4 (Reference 2.5-3, Subsection 2.4S.8.2.4) concludes that seiche oscillations in the MCR under the probable maximum hurricane (PMH) condition would be below the MCR embankment crest. According to Reference 2.5-3 (Subsection 2.4S.8.2.4), the PMH wind generates random waves with a spectral wave period of 4.7 seconds. It is concluded that the spectral period of such waves is significantly smaller than the natural period of the MCR which is approximately 22 minutes. Considering frictional effects, the energy of the oscillation will disperse and the raised water surface will decrease after each oscillation. In conclusion, the amplitude of the seiche oscillation will not be significant because the maximum wave height that can trigger a seiche motion is below the MCR embankment crest and there will be no flood risk to the plant as a result of seiche motions in the MCR.

2.5.4 Seiche in Other Water Bodies Seiche has not been considered the controlling mechanism for flooding in the Lower Colorado River, the Matagorda Bay and the Gulf of Mexico because the physical characteristics and locations of these water bodies are not conducive to the generation of significant seiche motions. Particularly, as described in Section 2.6 and in Subsection 2.5.2, historical tsunami database reports only one documented seiche event along the Texas Gulf Coast from the 1964 Alaska earthquake with a water height estimated to be on the order of less than one foot (based on gage record) to 6 feet (based on observers' account). Therefore, stream flooding is considered the primarily flooding mechanism on the Lower Colorado River and hurricane storm surge is the dominant factor responsible for coastal area flooding in the area. The flooding impact due to seiche effects in these offsite water bodies is considered insignificant and would be bounded by flooding induced by other mechanisms.

2.5.5 References 2.5-1 Office of Nuclear Regulatory Research, U.S. Nuclear Regulatory Commission, Design-Basis Flood Estimation for Site Characterization at Nuclear Power Plants in the United States of America. NUREG/CR-7046. November 2011.

2.5-2 American Nuclear Society, ANSI/ANS 2.8-1992, Determining Design Basis Flooding at Nuclear Power Reactor Sites, July 1992.

2.5-3 South Texas Project Units 3 & 4 Combined License Application (COLA), Final Safety Analysis Report (FSAR), Revision 7, Nuclear Innovation North America LLC (NINA),

February 1,2012.

2.5-4 STPEGS Updated Final Safety Analysis Report, Units 1 & 2; Section 9.2 and Subsection 2.5.6, Revision 16; Section 2.4, Revision 15.

2.5-5 U.S. Department of Commerce National Oceanic and Atmospheric Administration (NOAA), "Probable Maximum Precipitation Estimates, United States East of the 105th Meridian", Hydrometeorological Report No. 51, NOAA, June 1978.

2.5-6 U.S. Department of Commerce National Oceanic and Atmospheric Administration, "Application of Probable Maximum Precipitation Estimates - United States East of the 105th Meridian", Hydrometeorological Report No. 52, NOAA, August 1982.

Seiche Z.5-5

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushinma Response Project 2.5-7 U.S. Department of Commerce, "Two-to ten-day precipitations for return periods of 2 to 100 years in the contiguous United States", NWS Technical Paper 49, U.S, 1964.

2.5-8 U.S. Army Corps of Engineers, Coastal Engineering Manual, EM 1110-2-1100, Part II: Chapter 1, Chapter 2, Chapter 5 and Chapter 7, 1 August 2008; Chapter 3, 30 April 2002; Chapter 4, 31 July 2003; Part VI: Chapter 5, 28 Sep 11.

2.5-9 "Water wave Mechanics for Engineers and Scientists". Robert G. Dean and Robert.

Dalrymple, 1984, Prentice-Hall.

2.5-10 "Advanced Series on Ocean Engineering, Volume 15, Random Seas and Design of Maritime Structures", Y. Goda, 2000.

2.5-11 Not used.

2.5-12 "Historical Earthquakes, Prince William Sound, Alaska," United States Geological Survey, Available at http://earthquake.usgs.gov/regional/states/events/

1964_03_28.php, accessed April 27, 2007.

2.5-13 "Alaska Earthquake of 27 March 1964: Remote Seiche Stimulation. Science 145:

261-262," Donn, W. L. 1964.

2.5-14 "Seismic Seiches in Bays, Channels, and Estuaries, from The Great Alaska Earthquake of 1964: Oceanography and Coastal Engineering," McGarr, A. and R. C.

Vorhis. 1972, National Academy of Sciences, Washington, D.C. 25-28.

2.5-15 "Magnitudes and locations of the 1811-1812 New Madrid, Missouri, and the 1886 Charleston, South Carolina, Earthquakes," Bakun, W. H. and M. G. Hopper. 2004.

Bulletin of the Seismological Society of America 94(1): 64-75.

2.5-16 "Tsunamis and Tsunami-Like Waves of the Eastern United States," Science of Tsunami Hazards 20(3): 120-157, Lockridge, P. A., Lowell, S. W., and J. F. Lander.

2002.

Seiche 2.5-6

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Table 2.5-1 Short Duration PMP Depths at STP Site 6-hr,. 10 1-hr, Point PMP Depth PMP Duration & Area mi' Ratio Location Ratio Source (in) 72 hr, 10 mi2 HMR 51 - Fig. 22 55.7 48 hr. 10 mi2 HMR 51 -Fig. 21 51.8 24 hr, 10 mi2 HMR 51 - Fig. 20 47.1 12 hr, 10mi2 HMR_51_-Fig. 19 37.8 6 hr, 10 mi2 HMR 51 - Fig. 18 32.0 3 hr Fitted from Figure 2.5-3a 29.7 2 hr Fitted from Figure 2.5-3a 26.6 1 hr, point location 0.62 - HMR 52 - Fig. 23 19.8 30 min, point - 0.73 HMR 52 - Fig. 38 14.5 15 min, point 0.50 HMR 52 - Fig. 37 9.9 5 min, point 0.32 HMR 52 - Fig. 36 6.4 Source: Reference 2.5-3, Subsection 2.4S.2 Seiche 2.5-7

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Figure 2.5-1 ECP Embankment Cross-section 0 250 FeC Figure 2.5-2 STP I & 2 ECP Fetches and ECWIS Seiche 2.5-8

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project 60 A From HMR-51

-, 50

From HMR-52 M Interpolated for 2 hr Storm

.2 40 0 Interpolated for 3 hr Storm U - Log Fit of Data (HW-51 &H-R-52) 0 A

1 30 I S20 10 e" yy =ý0n2 7.6154Ln(x) + 21.31

= n wgi 0

0.01 0.1 1 10 100 Time Period (hr.)

Figure 2.5-3a PMP Depth versus Time Period for STP Site

From HMR-51

96-hour 10 square miles PMP (58.8 in)

Log Fit of Data 41 C

4-0.

0 D

0~

a- y = 9.5442In(x) + 15.277 R = 0.9925 x-Period (hr)

Figure 2.5-3b Determination of 96-hr PMP Depth for STP Site Seiche Z.5-9

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushinma Response Project 2.6 Tsunami Tsunami induced flood risk at the STP site was evaluated in support of the Combined License Application (COLA) for the future STP Units 3 & 4 (Reference 2.6-49). The tsunami flooding hazard reevaluation for STP 1 & 2 adopts the same approach and methodology as the COLA evaluation which was prepared in 2008, and supplement with recent data from literature and databases. The COLA evaluation followed the hierarchical approach described in NUREG/CR 6966 (Reference 2.6-1). Evaluation of tsunami hazard, as defined by Reference 2.6-1, requires the use of best available scientific information to arrive at a set of scenarios reasonably expected to affect a nuclear power plant site. The hierarchical screening process is based on a series of stepwise, progressively more refined analyses that evaluate hazards resulting from a tsunami. The hierarchical tsunami hazard assessment applied to the STP site include regional screening and site screening of tsunamigenic sources, and detailed numerical model simulations of the postulated probable maximum tsunami (PMT) generation scenarios and propagation in the Gulf of Mexico. In additional, the COLA evaluation is consistent with the requirement of NUREG/CR-7046 (Reference 2.6-50), which specifies that the antecedent water level of 10% exceedance high tide and the effect of sea level rise should be considered for tsunami flooding hazard assessment for coastal sites.

The STP 1 & 2 site is located about 3.2 mi west of the Lower Colorado River, and about 15 mi (24 kin), or 17 river miles, from the south Texas coast (Figure 2.6-1). Plant grade is at elevation 28 ft MSL (also referred to as NGVD 29). As demonstrated in the following, STP 1 & 2 will remain dry during the maximum tsunami flood level predicted for the area and the plant will not be exposed to any risk of tsunami flooding from the coast or from the river.

2.6.1 Probable Maximum Tsunami Tsunamis can be generated from near-field, mid-field or far-field sources. References 2.6-1, 2.6-2, 2.6-3, and 2.6-44 identify several types of tsunamigenic source mechanisms, including seismic events, volcanic events, submarine mass failures (SMFs), subaerial landslides, and impact of projectiles. With respect to tsunami hazard assessment for the Texas Gulf Coast near STP 1 & 2, the three primary forcing mechanisms are seismic events, volcanic events, and SMFs. These tsunami sources, in additional to an evaluation of local subaerial and submarine landside potentials adjacent to the Lower Colorado River near the STP site, are discussed in Subsection 2.6.3.

A hierarchical hazard assessment, based on an examination of historical records and characterization of possible tsunamigenic sources described in the following sections, concluded that the PMT for the STP 1 & 2 site is most likely to occur from an SMF similar to the East Breaks slump. Characterization and analysis of the East Breaks slump are discussed in detail in Subsection 2.6.4. However, as the interpretation of a single wave height from a slump scar may not be sufficient to bound the PMT flood risks on STP 1 & 2 due to the uncertainties inherent in the assessment, a range of hypothetical conditions were simulated at the East Breaks slump location. Simulations were performed using a hydrodynamic code known as the Method of Splitting Tsunami (MOST) (References 2.6-8 and 2.6-9). These simulations were intended to bracket any near-field tsunami hazard from a SMF event in the Gulf of Mexico.

Based on the results of these simulations, as described in Subsection 2.6.5, the peak flood level due to a PMT event is conservatively estimated to be on the order of 11.5 ft (3.52 m) MSL, which is below the STP 1 & 2 plant grade of 28 ft MSL and the flood level postulated from the Tsunami 2.6-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushimna Response Project Main Cooling Reservoir (MCR) breach event as described in Section 2.3. The PMT is therefore not the controlling flooding event for STP 1 & 2 safety-related systems, structures, and components (SSCs).

Subsequent to the 2008 tsunami evaluation for STP 3 & 4 COLA, a report by the U.S.

Geological Survey published in 2009 (Reference 2.6-44) on the regional assessment of tsunami potential in the Gulf of Mexico also developed numerical modeling simulations to evaluate tsunami flood levels in the region and concluded a maximum water level of 16.5 ft MSL, including 10% exceedance high tide and sea level rise, along the south Texas coast near the STP site. The USGS model used very conservative assumptions, more so than the STP 3 & 4 COLA analysis, especially with respect to the initial conditions that defined the water surface deformation at the hypothetical SMF sources. The predicted water levels are still well below both the plant grade elevation (28 ft MSL) and the controlling flood level for STP 1 & 2 from the breaching event of the MCR. The USGS model results are discussed further in Subsection 2.6.5.

2.6.2 Historical Tsunami Record 2.6.2.1 Tsunami Database Information and data on tsunami-generating earthquakes and runup events are included in the National Geophysical Data Center (NGDC) hazards database (References 2.6-18). The database contains information on source events and runup elevations for worldwide tsunamis from about 2000 BC to the present (Reference 2.6-1). Each event in the NGDC database has a validity rating used to indicate the probability that the event was a true tsunami. Similarly, each event includes a cause code identifying the forcing mechanism (e.g., earthquake, volcano, landslide, or any combination thereof). Information presented below from the NGDC database was confirmed in a more recent query of the tsunami database (Reference 2.6-46).

With respect to published literature, the publication titled "Caribbean Tsunamis: A 500 Year History from 1498-1998," is a compendium of data and anecdotal material on tsunamis reported in the Caribbean from 1498 to 1997 (Reference 2.6-19). Reference 2.6-20 includes source events and runup elevations for the Caribbean Sea and Eastern United States from 1668 to 1998, respectively. The USGS has published a fact sheet showing locations of plate boundaries in the Caribbean and tsunami-generating earthquakes from 1530 to 1991 (Reference 2.6-21).

The map is shown in Figure 2.6-2. Additionally, NOAA's Center for Tsunami Research, in conjunction with the Pacific Marine Environmental Laboratory, publishes information and analyses on tsunami sources and tsunami events (Reference 2.6-22).

The first documented tsunami event for the Texas coast occurred on October 24, 1918. This tsunami was reported to be caused by an aftershock of the Mw (Moment Magnitude) =7.5, October 11, 1918, earthquake near Puerto Rico (Reference 2.6-23, p. 73). As described in Reference 2.6-19 (p. 201), a small wave was recorded at the Galveston, Texas, tide gage. This event had a validity rating of four ("definite tsunami") in an earlier query of the database (Reference 2.6-18) but a more recent search indicated a validity rating of three ("probable tsunami") (Reference 2.6-46). The magnitude of tsunami runup was not reported.

The second documented tsunami event for the Texas coast occurred on May 2, 1922. The epicenter of the earthquake associated with this event was reported at 18.40 N and 64.9 0 W (Reference 2.6-19, p. 201). Reference 2.6-19 (p. 201) stated that "a wave with an amplitude of Tsunami 2.6-2

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushinma Response Project 64 cm was reported on a tide gage at Galveston. A train of three waves with a 45-minute period was followed in 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months by a 28-cm wave in a similar train of smaller waves. Parker [Reference 2.6-24] associated it with an earthquake felt 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months earlier at Vieques, Puerto Rico." However, according to Campbell (Reference 2.6-25, p. 56), the shock had a duration of only two seconds.

Therefore, the earthquake is unlikely to have been the tsunamigenic source. The validity rating of this event in the NGDC database is a two (i.e., a questionable event). No runups were documented along the Gulf coast for the primary shock of the 1922 earthquake. The surge was presumed to have been locally amplified by the inland position of the tidal gage (Reference 2.6-24, p. 30). The magnitude of the 1922 earthquake or the aftershock has not been estimated.

The third documented tsunami event for the Texas coast occurred on March 28, 1964 (Reference 2.6-46). The event was recorded on a tide gage in Freeport, Texas (Reference 2.6-26). While the validity code reported in the NGDC database is 4 ("definite tsunami") for this event, estimates of the wave height vary considerably between eyewitness accounts and tide gage data. Reference 2.6-26 (p. 261) notes that "inseveral reports from eyewitnesses in the coastal regions of Louisiana and Texas, waves up to 6 feet (2 meters) in height were observed."

However, Reference 2.6-26 (p. 261) also reports that the "maximum height of the recorded seiche at 0400 GMT is about seven inches (18 cm)," and that the "true wave height may have been several feet ([i.e.,] about a meter)." This event coincided with the 1964 Alaska (Mw=9.2) earthquake located between the Aleutian Trench and the Aleutian Volcanic Arc (Reference 2.6-27). Additional analyses of tide gage records from the 1964 event report the maximum measured height of the low-frequency waves along the Texas coast from the Alaska earthquake ranged from 0.22 to 0.84 feet (Reference 2.6-28, p. 26).

2.6.2.2 Tsunami Deposits Geologic criteria for the identification of tsunami deposits are discussed in Reference 2.6-2. As noted in Reference 2.6-2, a "combination of both the facies and sedimentology approach has resulted in an often-used, if not universally approved, set of criteria for understanding how sandy tsunami deposits might be distinguished in the stratigraphic record." These criteria include sand layers that are less than 25 cm thick and laterally continuous for hundreds of meters; a generally landward thinning sand sheet; an isochronous sand layer that typically cuts across stratigraphy; sands that contain a heterogeneous collection of marine microfossils; sands that are massive or plane laminated; evidence of erosion; decreasing grain size landward and upward, with possible inverse grading; and relative abundance of marine geochemical tracers such as bromine.

The distance from the location of the East Breaks slump to STP site is approximately 142 km (88 mi). As the area spanning this distance was likely to have included inorganic (e.g., sands, pebbles, gravels) and organic (e.g., large trees) material, geologic evidence for a tsunami, if present, would likely be composed of materials discussed in the preceding paragraph. No evidence of this type has been found.

The event deposits in Falls County, TX, have been interpreted by some investigators as being the result of a paleotsunami resulting from a large diameter (>10km) extraterrestrial bolide impact (Reference 2.6-52), while other investigators present alternative hypotheses, including sea level fall (Reference 2.6-53) and high energy storm events (Reference 2.6-54). Because of the uncertainty regarding the nature of the Falls County deposits, they are not considered in the evaluation of relevant tsunamigenic sources.

Tsunami 2.6-3

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project 2.6.3 Source Generator Characteristics Tsunamigenic source characteristics with potential to affect the US Atlantic and Gulf coasts are summarized in Reference 2.6-3, Reference 2.6-44, several databases, and published literature as discussed in the following subsections.

2.6.3.1 Seismic Tsunamis Sources of seismically induced tsunami events with the potential to impact the Gulf of Mexico were evaluated by References 2.6-3 and 2.6-4. Reference 2.6-3 (p. ii) stated that "tsunamis generated by earthquakes do not appear to impact the Gulf of Mexico coast." As stated in Reference 2.6-3 (p. 105):

"Earthquake-generated tsunamis generally originate by the sudden vertical movement of a large area of the seafloor during an earthquake. Such movement is generated by reverse faulting, most often in subduction zones. The Gulf of Mexico basin is devoid of subduction zones or potential sources of large reverse faults. However, the Caribbean basin contains two convergence zones whose rupture may affect the Gulf of Mexico, the North Panama Deformation Belt and the Northern South America Convergent Zone."

As stated in Reference 2.6-3, source areas with potential for tsunamigenesis affecting the US Gulf Coast include the North Panama Deformation Belt and the Northern South American Convergent Zone (Table 2.6-1). With respect to the North Panama Deformation Belt, Reference 2.6-3 stated that:

"the largest segment of the North Panama Deformation Belt is oriented between 60'-77'.

The 1882 Panama earthquake appears to have ruptured at least 3/4 of the available length of the convergence zone, and was estimated to have a magnitude of 8. While there was significant tsunami damage locally, there were no reports from the Gulf of Mexico of a tsunami from this earthquake. The low convergent rate (7-11 mm/yr) across the North Panama Deformation Belt supports long recurrence interval for large earthquakes."

The Global Centroid-Moment-Tensor (CMT) catalog was searched for potential seismogenic earthquakes in the two source regions of Table 2.6-1 (Reference 2.6-30). The following criteria were used for searching the CMT catalog within the North Panama Deformation Belt: a date range of January 1, 1976 (i.e., the start of the database) through November 4, 2008; latitude from 90 N to 120 N; longitude from 830 W to 770 W; depth from 0 to 1000 km; and moment magnitude (Mw) range from 6.5 to 10. The selection of a lower bound of Mw=6.5 is based on criteria from Reference 2.6-2 (p. 23) for a threshold moment magnitude of tsunamigenesis from earthquakes. One record was identified in the CMT catalog with these criteria. On April 22, 1991, a Mw=7.6 earthquake occurred at depth of 15 km and at a latitude of 10.100 N and a longitude of 82.770 W, located about 20 mi (32 km) offshore of the town of Limon, Costa Rica.

Source parameters for the earthquake were documented as a strike of 103 degrees, a dip of 25 degrees, and a rake of 58 degrees. Source parameters for earthquakes in the North Panama Deformation Belt with moment magnitudes below 6.5 are discussed in Reference 2.6-3. With respect to the far-field tsunami hazard on the south Texas coast, these additional sources are not reasonably expected to exceed the tsunamigenic potential of scenarios simulated by Reference 2.6-3 and Reference 2.6-4.

Tsunami 2.6-4

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushitna Response Project The following criteria were used for searching the CMT catalog within the Northern South American Convergent Zone: a date range of January 1, 1976, to November 4, 2008,; latitude from 11.50 N to 140 N; longitude from 770 W to 64' W; depth from 0 to 1000 km; and moment magnitude range from 6.5 to 10. No records were identified in the CMT catalog with these criteria. By broadening the criteria to include earthquakes from 0<Mw<10, two earthquake records were returned. The moment magnitude of these two earthquakes was 5.1, which is below the generally accepted threshold required for seismic tsunamigenesis as defined by Reference 2.6-2 (p. 23). A subsequent search of the CMT catalog in 2012 (Reference 2.6-45) using an extended date range (January 1, 1976 to November 6, 2012) and the same query parameters described above did not reveal any additional events.

Therefore, the assessment of far-field tsunami hazards in this region was based on tsunami simulations in References 2.6-4 and 2.6-3. Reference 2.6-4 performed tsunami simulations of seismic-borne tsunamis from postulated "worst case" events using a two-dimensional depth-integrated hydrodynamic model described in Reference 2.6-31. The following cases were used in the assessment (Reference 2.6-4, p. 305):

1. Mw = 9.0 at 66' W and 180 N (Puerto Rico trench);

2. Mw = 8.2 at 850 W and 210 N (Caribbean Sea);

3. Mw= 9.0 at660Wand 120 N; and

4. M, = 8.2 at 950 W and 200 N (near Veracruz, Mexico).

The source location of Case 3 at 660 W and 120 N is cited in Reference 2.6-4 (p. 305) as the North Panama Deformation Belt, but the location corresponding to 660 W and 120 N is the South Caribbean Deformed Belt (Reference 2.6-3, p. 110).

Source parameters for the model cases in Reference 2.6-4 were based on the formulae of Reference 2.6-32. For example, source parameters for the Veracruz scenario (Reference 2.6-4,

p. 305) are provided in Table 2.6-2. Reference 2.6-4 (p. 305) stated that the model sources were aligned with local strike.

Reference 2.6-4 (p. 311) concluded that "sources outside the Gulf are not expected to create a tsunami threatening to the Gulf coast." Reference 2.6-4 attributed this result primarily due to friction losses as the waves travel through the Straits of Florida and throughout islands in the Caribbean. Tsunami simulations in Reference 2.6-3 complemented earlier work by Reference 2.6-4, with Reference 2.6-3 (p. 117) stating that:

"in general, these results are consistent with the findings of Knight (2006)

[Reference 2.6-4], where the far-field tsunamis generated from earthquakes located beneath the Caribbean Sea are higher along the Gulf coast than the Atlantic coast because of dissipation through the Greater Antilles islands.

Conversely, tsunamis generated from earthquakes north of the Greater Antilles are higher along the Atlantic coast than the Gulf coast."

Reference 2.6-4 (p. 311) stated that one reason for this conclusion was that "the Atlantic and Gulf coasts are nearly independent since the hydrodynamic connection between basins is Tsunani 2.6-5

Enclosure NOC-AE-13002975 FloodingHazardReevaluation Report STP I & 2 Fukushima Response Project through the narrow Straits of Florida and through the Caribbean, where bottom friction losses appear to be large."

Additionally, the largest deepwater wave from the Reference 2.6-3 simulations was produced from the north Venezuela subduction zone. The maximum wave height from the north Venezuela subduction zone from a buoy at a depth of 250 m offshore of New Orleans, Louisiana, was estimated to be 6 cm (Reference 2.6-3, p. 130, Figure 74e, "Station 1").

While tsunamigenic earthquakes within the Gulf of Mexico have not been recorded, Reference 2.6-4 included a tsunami simulation assuming a magnitude Mw=8.2 earthquake offshore of Veracruz, Mexico. The resulting wave amplitude at the south Texas coast was about 0.35 m.

Intraplate earthquakes are less common than earthquakes occurring on faults near plate boundaries, but several earthquakes in the past three decades had epicenters within the Mississippi Canyon and Fan province (Reference 2.6-3). In recent time, the most severe earthquake in this region occurred on September 10, 2006. The moment magnitude was recorded as Mw = 5.8. The United States Geological Survey (USGS) concluded that earthquakes of this magnitude are unlikely to produce a destructive tsunami (Reference 2.6-33).

The second largest earthquake in this region occurred on February 10, 2006 with a moment magnitude of M, = 5.2. Reference 2.6-44 concluded that:

"There are no significant earthquake sources within the Gulf of Mexico that are likely to generate tsunamis, despite recent seismic activity in the area. Tsunami propagation from significant earthquake sources outside the Gulf of Mexico, such as the northern Panama Convergence Zone, Northern South America, Cayman Trough, the Puerto Rico trench, or the Gibraltar area shows that wave amplitude is greatly attenuated by the narrow and shallow passages into the gulf, and as a result, these tsunami sources do not constitute a tsunami hazard to the Gulf of Mexico coast."

Therefore, it is improbable that the PMT would be generated by a seismic event unless the seismic event induced a submarine landslide. That scenario is evaluated in Subsection 2.6.3.4.

2.6.3.2 Seismic Seiches Seismic seiches and the potential impact on STP 1 & 2 are described in Section 2.5.

2.6.3.3 Volcanism-based Tsunamis Reference 2.6-3 did not cite a tsunami hazard to the Gulf coast from volcanism. For example, Reference 2.6-3 stated that "far-field landslides, such as in the Canary Islands, are not expected to cause a devastating tsunami along the U.S. Atlantic coast." Previous studies have conjectured that the eruption and collapse of the Cumbre Vieja volcano on the island of La Palma in the Canary Islands could potentially affect the coast of Florida, USA, with a 25 m wave (Reference 2.6-5). A recent assessment of Reference 2.6-5 was discussed in Reference 2.6-3 (p. 57):

"as envisioned by Ward and Day (2001) [Reference 2.6-5], a flank collapse of the volcano may drop a rock volume of up to 500 km 3 into the surrounding ocean. The ensuing submarine slide, which was assumed to propagate at a speed of 100 m/s, will generate a strong tsunami with amplitudes of 25 m in Florida. In addition,

[Ward and Day, 2001] claimed that the collapse of Cumbre Vieja is imminent. In Tsunami 2.6-6

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STPI & 2 Fukushima Response Project our opinion, the danger to the U.S. Atlantic coast from the possible collapse of Cumbre Vieja is exaggerated. Mader (2001) [Reference 2.6-35] pointed out that Ward and Day's (2001) assumption of linear propagation of shallow water waves is incorrect, because it only describes the geometrical spreading of the wave and neglects dispersion effects. A more rigorous hydrodynamic modeling by Gisler et al. (2006) [Reference 2.6-36], confirms Mader's criticism. Their simulations show significant wave dispersion and predict amplitude decay proportional to r1 for a 3-dimensional model and r- 1 85 for a 2-D model (r is distance). [Reference 2.6-36]

predicted [a] wave amplitude for Florida is between 1 [and] 77 cm. [Reference 2.6-36 used] slightly smaller volume, 375 km 3, than Ward and Day (2001), but a much higher slide speed, that is much closer to the phase speed for tsunamis in the deep ocean (4,000 m of water)."

Further research on the La Palma event indicated that the distribution of slide blocks on the ocean bottom suggests that the collapse of Cumbre Vieja may not have been the result of a single catastrophic event, but the result of several smaller events. A recent report on potential tsunami threats to the United Kingdom concluded that "studies of the offshore turbidities [i.e.,

poorly sorted sediment that is deposited from a density flow of mixed water and sediment]

created by landslides from the flanks of the Canary Islands suggest that these result from multiple landslides spread over periods of several days" and are therefore "likely to create tsunamis of only local concern" (Reference 2.6-37, p. 23 and p. 30, respectively).

The National Geophysical Data Center (NGDC) natural hazard database for volcanoes (Reference 2.6-51) lists only two volcanoes (Los Atlixcos and San Martin) within 16 km (10 mi) of the present day Gulf of Mexico shoreline. Both volcanoes are located near Veracruz, Mexico.

Los Atlixcos is located about 9 km (5.6 mi) from the shoreline, while San Martin is located about 13 km (8.0 mi) from the shoreline. Based on the distance to the shoreline and proximity to the site, volcanogenic sources near Veracruz, Mexico are not expected to pose a flooding hazard to safety-related functions of the plant.

As no tsunamis have been documented in the Gulf of Mexico as a result of recent volcanic eruptions or associated mass wasting events (i.e., gravity-driven mass movement of soil, regolith, or rock moving downslope), this mechanism is not considered further as a potential source of tsunamis along the south Texas coast.

2.6.3.4 Submarine Slump Tsunamis Reference 2.6-3 (p. 35) cites four credible SMF source areas in the Gulf of Mexico: the Florida Escarpment, Campeche Escarpment, Northwest Gulf of Mexico, and the Mississippi Canyon (Figure 2.6-3). These four SMF source areas are located in three geologic provinces: a carbonate province, a salt province, and a canyon to deep-sea fan province. Multiple events have been identified for each scar. Many scars in these provinces correspond with relic events throughout the Quaternary (i.e., from 2.6 million to about 7500 years before the present, or yr BP). Notably, the geomorphology of SMFs in the Gulf of Mexico has been shown to be coupled with changes in sea level (Reference 2.6-6 and Reference 2.6-7). Reference 2.6-6 documents sea-level changes over the last 140,000 years, with the last lowstand of 120 m below present sea level occurring less than 20,000 years ago.

Tsunami 2.6-7

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP I & 2 Fukushima Response Project Carbonate Province The postulated SMF sources in the carbonate province are located offshore of West Florida and in the Campeche Escarpments north of the Yucatan Peninsula (Reference 2.6-3). The largest scar in this region is along the central part of the West Florida Slope and is estimated as 120 km long, 30 km wide, with a total volume of material removed of about 1,000 km3. However, formation of the scar was believed to have occurred as a result of multiple events. Most of the sediment was estimated to have been removed before the middle of the Miocene [c. 11.6 million years ago]. Reference 2.6-3 (p. 28) stated the following:

"During the Mesozoic, an extensive reef system developed around much of the margin of the Gulf of Mexico Basin by the vertical growth of reefs and carbonate shelf edge banks. This reef system is exposed along the Florida Escarpment and the Campeche Escarpment that fringe the eastern and southern margins of this basin. These escarpments stand as much as 1,500 m above the abyssal plain floor, and have average gradients that commonly exceed 200 and locally are vertical. Reef growth ended during the Middle Cretaceous, and subsequently the platform edges have been sculpted and steepened by a variety of erosional processes."

Salt Province The salt province is located in the northwestern Gulf of Mexico. Reference 2.6-3 (p. 32) stated that Geologic Long-Range Inclined Asdic (GLORIA) imagery identified 37 SMFs in the salt province and along the base of the Sigsbee Escarpment. The largest of these landslides is the East Breaks slump. With respect to the morphology of the salt province, Reference 2.6-3 (pp.

27-28) stated the following:

"Salt deposited in the late Jurassic Gulf of Mexico basin, the Louann salt, originally underlay large parts of Louisiana, southern Texas, and the area offshore of Mexico in the Bay of Campeche. As sediment eroded from the North American continent was deposited on this salt sheet throughout the Mesozoic and Cenozoic, the increased load caused the salt to flow with it migrating southward from the source area into the northern Gulf of Mexico. Presently the Louann salt underlies large parts of the northern Gulf of Mexico continental shelf and continental slope. South of Louisiana and Texas, the Sigsbee Escarpment is a pronounced cliff that marks the seaward limit of the shallowest salt tongue. As the salt is loaded, it flows both seaward and also upward through the overlying sediment column as cylindrical salt domes. The morphology of the salt sheet varies considerably across the margin. Salt domes are most common under the continental shelf, and most of the original salt sheet between individual domes in this region has been removed in response to the sediment loading, and migrated farther seaward."

The East Breaks slump is located approximately 88.2 mi (142 km) to the southeast of the STP 1

& 2 site (Figure 2.6-4), and its coordinates are approximately 27.57' N and 95.640 W. The slump is comprised of an eastern lobe and a western lobe. Reference 2.6-38 (p. 2) states that "the western and eastern lobes are thought to have formed by two different processes, and actually at two different, but relatively close, time periods. The western lobe formed as slump and debris deposits traveled downslope. The eastern lobe is more consistent with turbidity flow Tsunami 2.6-8

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 Fukushimna Response Project currents in the upper parts of the slide and leveed channels in the middle and lower portions of the slide." Further, Reference 2.6-38 (p. 3) states that "the eastern lobe appears more channelized and consists of density flow-type fill with few large slump and intact blocks. The western lobe, therefore, carried the bulk of the failed material and the energy level of the failure was much greater." As the eastern lobe was unlikely to have influenced tsunamigenesis, only the western lobe was used for the simulations.

The age of the East Breaks slump is not precisely known. Reference 2.6-39 (p. 366) states that the most recent mass wasting event responsible for the formation of the western lobe occurred about 16,000 yr BP and after the formation of the bulk of the eastern lobe. Reference 2.6-7 states that "the East Breaks Slide is a site of [sea level] lowstand instability, and seismic

[reflection] data shows repeated slope failure in this area. During late Quaternary lowstands of sea level, large deltas built up along the Texas-Louisiana shelf margin, and the present continental shelf [became] exposed as a subaerial coastal plain." Reference 2.6-7 also states that "itis clear that most sliding on the Texas-Louisiana slope occurred during the late Pleistocene [c. 10,000 -29,000 years BP] lowstands of sea level when sedimentation rates on the upper slope were high."

With respect to stability, Reference 2.6-3 notes that information on the age of landslides in the salt province is limited. Most landslides appear to have been active during oxygen isotope stages 2, 3, and 4 (18,170-71,000 yr BP) when salt movement due to sediment loading was most active. The age of the most recent landslide is less well established. Reference 2.6-7 states that that no major SMFs have occurred in the northwestern Gulf of Mexico in the Holocene (i.e., the last 10,000 years) and also states (on p. 309):

"Studies of submarine slides invariable prompt the question: Is the slope now completely stabilized? It is clear that most sliding on the Texas-Louisiana slope occurred during the late Pleistocene lowstands of sea level when sedimentation rates on the upper slope were high. No major Holocene slides have been documented. Low rates of deposition may be a primary reason for the present stability over much of the upper slope, and a further indication that sediments are relatively stable."

However, Reference 2.6-3 suggests the occurrence of at least one landslide during the Holocene, with "one unpublished age date of a sample below a thin landslide deposit (<3 m thick) indicates that it is younger than 6,360 yr BP." Therefore, no major SMFs have been documented for the salt province in over 6,300 years.

With respect to the dimensions of the East Breaks slump scar, estimates of width, length, area, and volume have varied with different studies. For example, Reference 2.6-40 states that the slump "consists of a 20-km wide head scarp initiated along the 150-meter isobath, a 55-km long erosional chute, ending in a 95 km by 30 km accretionary lobe. Total extent of the feature is 160 2 km from the shelf edge to a depth of 1,500 m" and "slumped deposits extend over a 3,200-km area with a volume on the order of 50-60 km3.'"Reference 2.6-7 states that "the East Breaks Slide is a prominent mass-transport feature. Revised bathymetry shows that the slide originated on the upper slope (200-1000 m), in front of a sandy late Wisconsinan shelf-margin delta, where the gradient is up to 30. It was deposited in a middle slope position (1000-1500 m) where the gradient is about 0.5'. Side-scan sonar data indicates that the slide is a strongly backscattering feature extending more than 110 km downslope from the shelf edge." Reference 2.6-3 (p. 32) states that "the largest of these failures occurs in the northwestern Gulf of Mexico, is 114 km Tsunani 2.6-9

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project long, 53 km wide, covers about 2,250 kiM2 , and has been interpreted to consist of at least two debris flows."

Source parameters for the East Breaks slump were estimated using three arc-second bathymetry data from the National Geophysical Data Center (NGDC) (Reference 2.6-41).

Source parameters, including slump width, length, and thickness, were estimated using a Geographic Information Systems (GIS) environment (Figure 2.6-5). Slump width was estimated to be approximately 13.4 km. The length of the erosional chute was estimated to about 42 km.

Based on a transect across the erosional chute, slump thickness was estimated to be about 100 m (see Path Profile A to A' in Figure 2.6-5). With respect to slope, Reference 2.6-40 states that "initial failure of the slump took place on very low angle slopes of less than two degrees while present slump deposits have an average seafloor slope of one-degree." While a vertical drop of 850 m over a length of 42 km indicates a bed slope of approximately 1.10, local bed slopes measured in GIS using a longitudinal transect along the erosional chute indicate a local maximum slope of about 1.950. Therefore, a maximum local slope of 20 was used for a conservative estimate. Similarly, initial depth of the slide was estimated conservatively using the 200-m and 1000-m bathymetry contour elevations. Therefore, initial depth was estimated to be 600 m (i.e., (200 m + 1000 m)/2) (Figure 2.6-5). Total length of the slide was taken from Reference 2.6-3 as 114 km.

Other SMFs identified in the salt province have areas that are an order of magnitude lower than the East Breaks slump (Reference 2.6-3) and are not further considered.

Canyon to Deep-sea Fan Province Three canyon to-deep-sea fan systems were formed during the Pliocene and Pleistocene: the Mississippi, Eastern Mississippi, and Bryant systems (Figure 2.6-3). The Mississippi system is the largest of the three systems. Borings and seismic data from the head of Mississippi Canyon indicate that there were alternating episodes of canyon filling and excavation between 19,000 and 7,500 years before the present (YBP). Also, Geologic Long-Range Inclined Asdic (GLORIA) imagery of the Mississippi Fan suggests that this feature consists of at least two separate events (Reference 2.6-3). According to Reference 2.6-3, the resumption of hemipelagic sedimentation at the head of the Mississippi Canyon by 7500 yr BP indicates that the largest of the landslide complexes ceased being active by the middle of the Holocene. The largest SMF in the complex covers approximately 23,000 km 2 and reaches 100 m in thickness, with a volume estimated to be about 1,750 km3 .

As with the East Breaks slump, estimates of the maximum credible landslide scar dimensions for the Mississippi Fan have varied with different investigations. For the maximum credible single event, Reference 2.6-48 reported a volume of 426 kiM3, with a corresponding area of 3,687 km 2. Reference 2.6-3 and Reference 2.6-48 cited the excavation depth as 300 m in the upper canyon, with a runout length of 442 km from the headwall scarp.

The Eastern Mississippi and Bryan Canyon systems are smaller than the Mississippi Canyon system. The Eastern Mississippi system has a deposit that is "approximately 154 km long, as much as 22 km wide, and covers an area of 2,410 km 2"' (Reference 2.6-3, p. 34). With respect to the Bryant system, Reference 2.6-3 (pp. 33-34) states that "The Bryant Canyon system was immediately downslope of a shelf edge delta system, and failure of this system has been proposed as the explanation for Tsunami 2.6-10

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPI & 2 Fukushima Response Project thick chaotic deposits in mini basins along the path of this canyon system Debris from the failure of the shelf edge delta was transported down the Bryant Canyon system, but these landslide deposits predate and are buried by the smaller landslides off the mini-basin walls."

Based on the above information, an SMF is considered to be the bounding tsunami generator in the Gulf of Mexico. Simulation based on postulated SMF events were performed to estimate the PMT water levels at the STP & 2 site, as described in Subsection 2.6.4 below.

2.6.3.5 Onshore Landslides The potential for large scale subaerial or submarine landslides adjacent to the Lower Colorado River, which is relatively shallow in the vicinity of the STP site, is highly improbable due to the flat terrain. This is consistent with the conclusion of the UFSAR for STP 1 & 2 (Reference 2.6-47). Therefore, based on the topography near the site and the shallow bathymetry of the river, it is evident that there will not be any threat of flooding posed to STP 1 &2 due to surges from bank material sliding into the Lower Colorado River.

2.6.4 Tsunami Analysis Tsunami modeling was conducted for a hypothetical SMF event originating at the location of the East Breaks slump near the south Texas coast. Consistent with the approach in Reference 2.6-12 and Reference 2.6-13, a series of scaled dipolar initial conditions were used for bracketing a conservative range of initial wave heights. Hydrodynamic simulations were modeled using a series of codes collectively known as the Method of Splitting Tsunami (MOST) (References 2.6-8), which has been subject to extensive validation testing (Reference 2.6-2 and Reference 2.6-9). For all model simulations, maximum runup along the south Texas coast did not exceed 2 m (6.56 ft) above Mean Sea Level (MSL).

MOST is formulated based on the three phases of long wave evolution (Reference 2.6-8):

(i.) A "Deformation Phase" that generates the initial conditions for a tsunami by simulating ocean floor and corresponding free surface changes due to a forcing mechanism; (ii.) A "Propagation Phase" that propagates the generated tsunami across the deep ocean using Nonlinear Shallow Water (NSW) wave equations; and (iii.) An "Inundation Phase" that simulates the shallow ocean behavior of a tsunami by extending the NSW calculations using a multi-grid runup algorithm to predict coastal flooding and inundation.

The three-dimensional (two horizontal and one vertical) NSW equations are solved numerically using a finite difference algorithm that splits the NSW equations into a pair of systems and a series of nested grids, which are discussed further below.

For wave generation, initial wave dimensions (initial elevation of the depression wave due to a slump) were estimated using the slump center of mass motion model which is based on curve fits from sliding block experiments described in Reference 2.6-10 and Reference 2.6-11 b.

Source parameters documented in the paragraphs above and in Figure 2.6-5 were used for Tsunami 2.6-11

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 Fukushima Response Project estimating initial wave height. Specific gravity of the slump mass was assumed to be equal to 2 (Reference 2.6-1 Ia). The 100-m thickness (T) with respect to the 600-m initial depth (h)

(T/h=0.17) and the thickness relative to the 42 km length (b) of the erosional chute (T/b=0.002) suggest that the initial wave height from the East Breaks slump would be relatively small. Using the NGDC bathymetry data (Figure 2.6-5), initial wave height for the East Breaks slump was estimated to be 7.9 m. Considering the variability in interpreting landslide dimensions, the estimated wave height of 7.9 m is similar to the tsunami wave height predicted by Reference 2.6-40 (on the order of 7.6 meters).

Many previous SMF tsunami studies have assumed simplified wave shapes for the initial tsunami wave (Reference 2.4.6-14). For the MOST simulations, specification of an initial deformation condition (horizontal area of sea surface deformation from 0 ft MSL due to initial wave) was based on scaling a dipole wave. A dipole wave is similar to the structure of an N-wave, which is a wave with a leading negative or depression wave followed by a positive elevation wave. Scaling of the wave dimensions into a dipole condition is based on information from other SMF events (Reference 2.4.6-12 and Reference 2.4.6-13) and estimated source parameters for the East Breaks slump from Subsection 2.6.3.

SMF events used for initial wave dimensions include the Palos Verdes (PV) landslide in Southern California (Reference 2.6-12) and the 1998 Papua New Guinea (PNG) slump in the Sandaun Province (Reference 2.6-23). Initial conditions from other events were used as relatively little data exists for SMF tsunamis, and the PV and PNG events have been analyzed extensively by the tsunami community (Reference 2.4.6-14). The PV case was used as a "lower bound" or base case condition. An upper bound condition was developed by assuming an almost instantaneous characteristic time for the SMF. The "upper bound" case was based on the PV case and scaled up by twenty times (PV20). This condition was used to set a reasonable upper limit of wave height for the East Breaks slump. A hypothetical "Monster" condition (hereinafter referred to as "Monster") was also developed for the East Breaks slump to test a very wide initial wave, as opposed to only a tall and steep initial wave. The "Monster" condition has not been simulated or described previously in the tsunami literature.

Initial deformation areas for each simulation (PV, PV20, PNG, and "Monster") are provided in Table 2.6-3. Initial deformation areas range from about 387 km 2 to about 9,932 km 2 (149 mi 2 to 3835 mi 2, respectively).

After specifying an initial deformation condition, the propagation phase is based on a simplified form of the Navier-Stokes equations referred to as the nonlinear shallow water (NSW) equations (Reference 2.6-8). The NSW equations are solved numerically with a finite difference algorithm and a series of nested grids (Reference 2.6-42). Since tsunami wavelength becomes shorter during shoaling, a series of nested grids are required for maintaining resolution of the wave with decreasing water depth. Therefore, three grids (i.e., A, B and C) were used for the MOST simulations (Figure 2.6-6). The grids were derived from NGDC topography and bathymetry data (Reference 2.6-41). Grid spacing between cell nodes was equal to 12 arc-seconds (about 360 m), 6 arc-seconds (about 180 m), and 6 arc-seconds (about 180 m), respectively.

For the inundation phase, MOST uses a moving boundary calculation for estimating tsunami runup onto dry land. Details of the moving boundary are discussed in References 2.6-8, 2.6-14 and 2.6-42. While friction factors are not used in the propagation phase of MOST, one must be specified for the inundation phase. Reference 2.6-2 states that "several studies show that an unsteady flow during runup is not very sensitive to changes in the roughness coefficient", and Tsunami 2.6-12

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushimna Response Project that "any moving boundary computation induces numerical friction near the tip of the climbing wave (except in a Lagrangian formulation)." Sensitivity tests for the MOST inundation friction factor were performed with the PNG simulation case. The inundation friction factor was set equal to a value corresponding to Manning's n=0.032. In addition to Manning's n=0.032 that was used in the MOST simulations, values of Manning's n=0.010, Manning's n=0.011 and Manning's n=0.016 were also tested for impacts to the results. Lower values of the Manning's coefficient have a negligible impact on flooding near the STP site since simulated tsunami inundation is limited primarily to the barrier islands to the east of the barrier island.

Initial wave dimensions for the PV, PV20, PNG, and "Monster" simulation cases are shown in Figures 2.6-7 to 2.6-14. Initial wave elevations are relative to a still water level of 0 m (0 ft) MSL.

Initial wave trough elevations varied from -7 m to -140 m (-23 ft to -459 ft, respectively) (MSL) and initial wave crest elevations varied from 3 m to 60 m (9.8 ft to 197 ft, respectively) (MSL)

(Table 2.6-3). Initial wave widths varied from 14 km (8.7 mi) to 136 km (85 mi).The range of initial wave heights and initial wave widths were intended to simulate reasonably probable bounding cases for SMFs that may occur offshore of the South Texas coast.

2.6.5 Tsunami Water Levels 2.6.5.1 MOST Modeling Results As far-field tsunamis are unlikely to impact the south Texas coast, the PMT is simulated as a tsunami generated by a near-field submarine landslide near the East Breaks slump. Using the MOST code (Reference 2.6-8), a series of scaled initial conditions were used to assess the near-field hazard of tsunami generation from submarine landslides to the STP I & 2 site. Four scenarios with wave heights ranging from -140 m (-459 ft) to 60 m (197 ft) and deformation areas ranging from 387 km 2 to 9932 km 2 (Table 2.6-3) were simulated. The model results indicate that tsunami waves from the SMFs diffuse rapidly by the continental shelf offshore of the south Texas coast. The remaining wave energy that reached the south Texas coast was largely reflected by the barrier islands, and thus the maximum predicted runup from the simulations did not exceed 2 m. Maximum flow depth from the simulations, which occurred at the shoreline, did not exceed 3.25 m. Maximum rundown did not exceed 2.5 m about 1 mi offshore.

Relative to the location of the STP 1 & 2 site, most SMF sources in the Gulf of Mexico are mid-field to far-field sources, i.e., source locations over 200 km away (Figure 2.6-3). The distance from STP 1 & 2 to the East Breaks slump is 142 km (88.2 mi). The distance from STP 1 & 2 to Bryant Canyon is 517 km (321.2 mi). The distance from STP 1 & 2 to Mississippi Canyon and the Eastern Mississippi Canyon/Fan is 640 km (397.7 mi) and 709 km (440.6 mi), respectively.

The distance from STP 1 & 2 to the Campeche Escarpment and Bay of Campeche is 873 km (542.5 mi) and 953 km (592.2 mi), respectively. The distance from STP 1 & 2 to the Florida escarpment is 1169 km (726.4 mi). Since landslide waves tend to be steep (i.e., high initial wave height relative to wavelength) and are prone to breaking, wave heights at the East Breaks slump from mid-field and far-field sources are not expected to exceed the simulated initial conditions.

As shown with the simulations, diffusion and energy dissipation from large SMF events is likely to be significant. Therefore, potential runup from these events is likely to be lower than the scenarios modeled for the East Breaks slump, and additional landslide scenarios in the Gulf of Mexico are not considered further.

Tsunami 2.6-13

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STPJ & 2 Fukushimna Response Project MOST output includes maximum runup estimates (i.e., maximum inland elevation inundated by the tsunami above MSL). Maximum runup ranges from 1 to 2 m (3.28 to 6.56 ft, respectively)

MSL for the south Texas coast near the STP 1 & 2 site (Table 2.6-3). The simulations indicate that a landslide tsunami originating from the East Breaks slump location would be unlikely to cross the barrier islands and produce a runup in excess of 2 m (6.56 ft) MSL. Plots of maximum wave amplitude relative to south Texas coast bathymetry are shown for PV, PV(x20), PNG, and the hypothetical "Monster" condition in Figures 2.6-15, 2.6-17, 2.6-19, and 2.6-21, respectively.

Time series of wave amplitude for a buoy located near the south Texas coast for the PV, PV(x20), PNG, and hypothetical "Monster" are shown in Figures 2.6-16, 2.6-18, 2.6-20, and 2.6-22, respectively.

Estimation of the maximum flood level for a PMT event also included an analysis of the 10%

exceedance of the astronomical high tide and the consideration of the long-term sea level rise, per guidance in NUREG/CR-7046 (Reference 2.6-50). Based on tide gage data for NOS Station

8772440, the 10% exceedance of the astronomical high tide was estimated to be 3.54 ft MSL (Reference 2.6-16). The long-term sea level rise for this station was estimated by NOAA to be 1.43 ft per century (Reference 2.6-17). The peak flood level due to a probable maximum tsunami event is therefore estimated to be of the order of 11.5 ft MSL within the next century (i.e., 6.56 ft tsunami runup + 3.54 ft 10% exceedance of the astronomical high tide + 1.43 ft sea-level rise = 11.5 ft MSL).

With respect to the assumption of the MSL datum (or NGVD 29) shift relative to actual mean sea level from tidal measurements, it should be noted that the Freeport, Texas, tide gage does not have a published or official NGVD29 orthometric height mark. However, the one mark that does exist suggests that the difference between MSL to actual mean sea level is small (i.e.,

within +/-0.2 ft of the Mean Lower-Low Water datum), and thus 11.5 ft MSL should be considered as a reasonable approximation of the actual value.

Based on the discussion above it is concluded that the probable maximum tsunami event is not the controlling design basis flood event for the STP I & 2 site, because the postulated flood level is much lower than the design basis flood elevation predicted for a hypothetical breach event of the Main Cooling Reservoir embankment, described in Section 2.3. Coincident wind waves are not considered in the analysis since it is evident that the PMT event will not inundate the site and will have no flooding impacts on safety-related facilities of STP 1 & 2.

2.6.5.2 USGS Report (2009) - COULWAVE Modeling Results A recent assessment of tsunami potential in the Gulf of Mexico is provided in a 2009 USGS report (Reference 2.6-44). Of the four credible submarine mass failure (SMF) sources in the Gulf of Mexico (Northwest Gulf of Mexico, Mississippi Canyon, Florida Margin, and Campeche Margin), the 2009 USGS report suggests that the propagation paths that result in the least attenuation of potential tsunamis for the south Texas coast are for the East Breaks and Campeche provinces. The report (Reference 2.6-44) describes the hydrodynamic simulations that were performed for a potential source near the East Breaks slump and for two potential sources at the Campeche Escarpment (for a 20-kmn slide width and for a 60-km slide width). The numerical simulations were performed using the tsunami model COULWAVE, which solves the fully nonlinear extended Boussinesq equations on a Cartesian grid. The numerical scheme is based on a fourth order Adams-Bashforth-Moulton predictor-corrector time integration scheme, with spatial derivatives approximated with fourth order, centered finite differences.

Tsunami 2.6-14

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project With respect to the physical assumptions for the simulations, Reference 2.6-44 states that "the purpose of these initial simulations is to provide an absolute upper limit of the tsunami wave height that could be generated" and that "these limiting simulations use physical assumptions that are arguably unreasonable but provide maximum amplitude estimates." For the initial deformation of the water surface at the source, the time scale of the seafloor motion was assumed to be very small compared to the period of the generated water wave. With this assumption, the free water surface response matches the change in the seafloor profile exactly.

Therefore, for estimating the initial free surface condition, Reference 2.6-44 states that "the initial pre-landslide bathymetry profile, as estimated by examination of neighboring depth contours, is subtracted by the post (existing) landslide bathymetry profile. This difference surface is smoothed and then used directly as a "hot-start" initial free surface condition in the hydrodynamic model." Reference 2.6-44 notes that "this type of approximation is used commonly for subduction-earthquake-generated tsunamis, but is known to be very conservative for landslide tsunamis." As with the MOST simulations, the COULWAVE simulations assume initial wave elevations are relative to a still water level of 0 m (0 ft) MSL (also referred to as NGVD 29).

For tsunami propagation, the two horizontal dimension (2HD) COULWAVE simulations were based on a constant spatial grid size of 200 m. Also, bottom roughness was assumed to be negligible in areas that were initially wet (i.e., locations with negative bottom elevation).

Regarding inundation, the report states that "it is most reasonable to analyze the [COULWAVE]

2HD results only to the initial shoreline. The relatively coarse grid size used in the [COULWAVE]

2HD results might cause accuracy degradation during the inundation phase due to poor resolution of shallow bathymetric and on land features." The initial wave characteristics for the East Breaks slump tsunami in Reference 2.6-44 include an initial trough elevation assumed to be -160 m (-525 ft) (MSL) and an initial crest elevation assumed to be 100 m (328 ft) (MSL).

For the Campeche Escarpment (also referred to as Campeche Margin) tsunami assessment, the initial wave characteristics in Reference 2.6-44 are based on a 20-km slide width. The initial conditions were based on the maximum observed landslide for the Florida Escarpment due to the lack of detailed bathymetry data and distribution of landslides on or above the Campeche Margin. Reference 2.6-44 states that "as a provisional source for the Campeche Margin, we used initial conditions applicable to the maximum observed landslide along the slope above the Florida Margin, a similar geologic environment. This includes an initial drawdown of 150 m, with a horizontal length scale of 20 km." The initial trough elevation for the Campeche Escarpment tsunami was assumed to be -150 m (-492 ft) (MSL) and the initial crest elevation was assumed to be 150 m (492 ft) (MSL). Based on the results of synthetic tsunami time series (marigrams) for sources at the East Breaks slump and Campeche Escarpments at an ocean water depth of 50 m (MSL) near Matagorda Bay, Texas, Reference 2.6-44 identifies equal tsunami runup potential from the East Breaks slump and the Campeche Escarpment, stating the following:

"It was expected that because the propagation distance for Campeche is so much larger than East Breaks (about 700 km longer), the 2D spreading effect is significant, and results in greater attenuation than for the East Breaks scenario. Figure 4-11 compares the ocean surface elevation time series for the offshore Campeche 20-km wide slide and the East Breaks (2HD simulations) at the same 50-m depth offshore location. The general conclusion made from this comparison is that the approaching wave heights for the hypothetical Campeche scenario are comparable to that of the East Breaks scenario, unless it is found that the maximum slide width in the Campeche province is much less than 20 km. Because the properties of the incoming waves are different (leading Tsunami 2.6-15

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 Fukushima Response Project elevation vs. leading depression), and the uncertainty in the slide parameters, this analysis indicates that East Breaks and Campeche (20 km width) should have equal tsunami potential on the Texas coast."

Reference 2.6-44 cited "realistic wave propagation in two horizontal dimensions yielded potential maximum tsunami runup [along the Gulf Coast] of approximately 4 m mean sea level (MSL)." Therefore, as a result of independent hydrodynamic modeling using COULWAVE, Reference 2.6-44 states that "the potential maximum water level for the conservative 2HD tsunami over the next century is 4 m (maximum tsunami runup) + 0.45 m (10% exceedance high tide) + 0.59 m (century sea level rise) or approximately 5.0 m (16.5 ft) (MSL)." This maximum tsunami water level predicted by the COULWAVE model is lower than the STP 1 & 2 plant grade elevation of 28 ft MSL by more than 11 ft.

2.6.6 Hydrography and Harbor or Breakwater Influences on Tsunami Because the STP 1 & 2 site is over fifteen miles inland from the coast and barrier islands and the postulated maximum flood levels due to the PMT event is much lower than the STP 1 & 2 plant grade elevation of 28 ft MSL, there will be no local onsite adverse effects associated with tsunami events, including breaking waves, bores, or any resonance effects that would result in higher tsunami runup on the safety-related facilities.

2.6.7 Hydrostatic Forces, Hydrodynamic Forces, Debris, and Waterborne Projectiles Due to the location and elevation of the plant described in Subsection 2.6.6, flood waves generated by tsunamis are not expected to reach the STP 1 & 2 site. As such, any effects due to hydrostatic forces, hydrodynamic forces, debris, and waterborne projectiles are precluded.

2.6.8 Effects on Safety-Related Facilities The estimated maximum water level predicted by modeling using MOST is 11.5 ft MSL, while independent modeling performed using COULWAVE (Reference 2.6-44) predicts a maximum water level of 16.5 ft MSL under more conservative physical assumptions, most notably the "hot start" initial condition. Both estimates are lower than the plant grade elevation of 28 ft MSL.

Therefore, the PMT event will have no flooding impacts on safety-related facilities or the design basis functions of STP 1 & 2, and there will be no associated impact due to debris, water-borne projectiles, sedimentation including erosion and deposition on the safety-related facilities of STP 1 & 2.

Tsunami 2.6-16

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STP I & 2 Fukushimna Response Project 2.6.9 References 2.6-1 Prasad, R. and Pacific Northwest National Laboratory (PNNL), "Tsunami Hazard Assessment at Nuclear Power Plant Sites in the United States of America," NUREG CR-6966, PNNL-17397, Nuclear Regulatory Commission, Final Report, March 2009.

2.6-2 Gonzalez, F. I., E. Bernard, P. Dunbar, E. Geist, B. Jaffe, U. Kanoglu, J. Locat, H.

Mofjeld, A. Moore, C. E. Synolakis, V. Titov, and R. Weiss, et al. (Science Review Working Group), "Scientific and Technical Issues in Tsunami Hazard Assessment of Nuclear Power Plant Sites," National Oceanic and Atmospheric Administration (NOAA) Technical Memorandum OAR Pacific Marine Environmental Laboratory 136, 2007.

2.6-3 Atlantic and Gulf of Mexico Tsunami Hazard Assessment Group, "Evaluation of Tsunami Sources with the Potential to Impact the U.S. Atlantic and Gulf Coasts - A Report to the Nuclear Regulatory Commission," U.S. Geological Survey Administrative Report, August 22, 2008.

2.6-4 Knight, B., "Model Predictions of Gulf and Southern Atlantic Coast Tsunami Impacts from a Distribution of Sources," Science of Tsunami Hazards 24(2): 304-312, 2006.

2.6-5 Ward, S. N. and S. Day, "Cumbre Vieja Volcano - Potential Collapse and Tsunami at La Palma, Canary Islands," Geophysical Research Letters 28(17): 3397-3400, 2001.

2.6-6 Simms, A. R., J. B. Anderson, K.T. Milliken, Z. P. Taha, and Wellner, "Geomorphology and Age of the Oxygen Isotope Stage 2 (Last Lowstand) Sequence Boundary on the Northwestern Gulf of Mexico Continental Shelf," in Seismic Geomorphology: Applications to Hydrocarbon Exploration and Production, edited by R. J. Davies, H.W. Posamentier, L. J. Wood, and J. A. Cartwright, Geological Society, London, Special Publications, 277: 29-46, 2007.

2.6-7 Rothwell, R. G., N. H. Kenyon, and B. A. McGregor, "Sedimentary Features of the South Texas Continental Slope as Revealed by Side-Scan Sonar and High-Resolution Seismic Data," The American Association of Petroleum Geologists Bulletin 75(2): 298-312, 1991.

2.6-8 Titov, V. V., and F. I. Gonzalez, et al., "Implementation and testing of the Method of Splitting Tsunami (MOST) model," NOAA Technical Memorandum ERL PMEL1 12, Pacific Marine Environmental Laboratory, 1997.

2.6-9 Synolakis, C. E. and E. N. Bernard, "Tsunami Science Before and Beyond Boxing Day 2004," Philosophical Transactions of the Royal Society 64: 2231-2265, 2006.

2.6-10 Watts, P., "Tsunami Features of Solid Block Underwater Landslides," Journal of Waterway, Port, Coastal, and Ocean Engineering, 126(3): 144152, 2000.

2.6-1 la Grilli, S. T. and P. Watts, "Tsunami Generation by Submarine Mass Failure, Part I, Modeling Experimental Validation, and Sensitivity Analyses," Journal of Waterway, Port, Coastal, and Ocean Engineering, 131(6): 283-297, 2005.

Tsunami 2.6-17

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushiina Response Project 2.6-1lb Watts, P., S. T. Grilli, D. R. Tappin, and G. J. Fryer, "Tsunami Generation by Submarine Mass Failure, Part II, Predictive Equations and Case Studies," Journal of Waterway, Port, Coastal, and Ocean Engineering, 131(6): 298-310, 2005.

2.6-12 Borrero, J. C., Ph.D., "Tsunami Hazards in Southern California," Thesis, University of Southern California, 2002.

2.6-13 Synolakis, C. E., J. P. Bardet, J. C. Borrero, H. L. Davies, E. A. Okal, E. A. Silver, S.

Sweet, and D. R. Tappin, "The Slump Origin of the 1998 Papua New Guinea Tsunami," Proceedings: Mathematical, Physical and Engineering Sciences, 458(2020): 763-789, 2002.

2.6-14 Synolakis, C. E., "Tsunami and Seiche," in Earthquake Engineering Handbook, edited by W. F. Chen and C. Scawthorn, CRC Press, pp. 9-1 to 9-90, 2004.

2.6-15 Not Used.

2.6-16 National Oceanic and Atmospheric Administration (NOAA), "NOS Station #8772440, Freeport - Verified Historic Tide Data." Available at http://tidesandcurrents.noaa.gov/datamenu.shtml?stn=8772440%20Freeport, %20T X&type=Historic+Tide+Data, accessed August 19, 2008.

2.6-17 National Oceanic and Atmospheric Administration (NOAA), "NOS Station #8772440, Freeport - Sea Level Trends." Available at http://tidesandcurrents.noaa.gov/sltrends/sltrends-station.shtml?stnid=8772440%20 Freeport,%20TX, accessed August 19, 2008.

2.6-18 National Geophysical Data Center/World Data Center, "Tsunami Data at NGDC."

Available at http://www.ngdc.noaa.gov/hazard/tsu.shtml, accessed November 22, 2008.

2.6-19 0' Loughlin, K. F., and J. F. Lander, Caribbean Tsunamis: A 500-Year History from 1498-1999, Kluwer Academic Publishers, 280 pp., 2003.

2.6-20 Lockridge, P. A., S. W. Lowell, and J. F. Lander, "Tsunamis and Tsunami-Like Waves of the Eastern United States," Science of Tsunami Hazards, 20(3): 120-157, 2002.

2.6-21 United States Geological Survey, "Improving Earthquake and Tsunami Warnings for the Caribbean Sea, the Gulf of Mexico, and the Atlantic Coast," USGS Fact Sheet 2006-3012, 2006.

2.6-22 Pacific Marine Environmental Laboratory (PMEL), "NOAA Center for Tsunami Research." Available at http://nctr.pmel.noaa.gov, accessed November 23, 2008.

2.6-23 Lander, J. F., L. S. Whiteside, and P. A. Lockridge, "A Brief History of Tsunamis in the Caribbean Sea," Science of Tsunami Hazards, 20(2): 57-94, 2002.

2.6-24 Parker, W. E., "Unusual Tide Registration of Earthquake," Bulletin of the Seismological Society in America, pp. 28-30, 1922.

Tsunami 2.6-18

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 FukushirnaResponse Project 2.6-25 Campbell, J. B., "Earthquake History of Puerto Rico, Seismicity Investigation," An Earthquake History of Puerto Rico, Aguirre Nuclear Power Plant, Weston Geophysical Research," Appendix I, Part A, Weston, Massachusetts, 1972.

2.6-26 Donn, W. L., "Alaska Earthquake of 27 March 1964: Remote Seiche Stimulation,"

Science, 145: 261-262, 1964.

2.6-27 United States Geological Survey, "Historical Earthquakes, Prince William Sound, Alaska." Available at http://earthquake.usgs.gov/regional/states/events/

1964_03_28.php, accessed April 27, 2007.

2.6-28 McGarr, A., and R. C. Vorhis, "Seismic Seiches in Bays, Channels, and Estuaries,"

The Great Alaska Earthquake of 1964: Oceanography and Coastal Engineering, National Academy of Sciences, Washington, D.C. 25-28, 1972.

2.6-29 Not Used.

2.6-30 "Global Centroid Moment Tensor (CMT)," Available at http://www.globalcmt.orgl, accessed November 21, 2008.

2.6-31 Kowalik, Z., W. Knight, T. Logan, and P. Whitmore, "Numerical Modeling of the Global Tsunami: Indonesian Tsunami of 26 December 2004," Science of Tsunami Hazards, 23(1): 40-56, 2005.

2.6-32 Okada, Y., "Surface Deformation Due to Shear and Tensile Faults in a Half-Space,"

Bulletin of the Seismological Society of America, 75(4): 11351154, 1985.

2.6-33 National Earthquake Information Center, "Magnitude 5.8 Gulf of Mexico Earthquake of 10 September 2006," United States Geological Survey. Available at http://earthquake.usgs.gov/eqcenter/eqinthenews/2006/usslav/#summary, accessed December 02, 2008.

2.6-34 Not Used.

2.6-35 Mader, C. L., "Modeling the La Palma Landslide Tsunami," Science of Tsunami Hazards, 19(3): 150-170, 2001.

2.6-36 Gisler, G., R. Weaver, and M.L. Gittings, "SAGE Calculations of the Tsunami Threat from La Palma," Science of Tsunami Hazards, 24(4): 288-301, 2006.

2.6-37 British Geology Society, "The Threat Posed by the Tsunami to the UK," Study commissioned by Debra Flood Management, 133 pp., 2005.

2.6-38 Hoffman, J. S., M.J. Kaluza, R. Griffiths, J. Hall, and T. Nguyen, "Addressing the Challenges in the Placement of Seafloor Infrastructure on the East Breaks Slide-A Case Study: The Falcon Field (EB 579/623)," Northwestern Gulf of Mexico, Offshore Technology Conference 16748, 2004.

Tsunami 2.6-19

Enclosure NOC-AE-13002975 Flooding Hazard Reevaluation Report STP I & 2 Fukushimna Response Project 2.6-39 Piper, J. N., and Behrens, "Downslope Sediment Transport Processes and Sediment Distributions at the East Breaks, northwest Gulf of Mexico," Proceedings of the 23rd Annual Gulf Coast Section SEPM Research Conference, Houston, Texas, pp. 359-385, 2003.

2.6-40 Watts, P., F. Lettieri, and G. Jamieson. "East Breaks Slump, Northwest Gulf of Mexico," Offshore Technology Conference, OTC paper 12960, 2001.

2.6-41 National Geophysical Data Center (NGDC), "Bathymetry, Topography, and Relief,"

National Oceanic and Atmospheric Administration (NOAA). Available at http://www.ngdc.noaa.gov/mgg/bathymetry/relief.html, accessed October 5, 2008.

2.6-42 Titov, V. V., and C. E. Synolakis, "Numerical Modeling of Tidal Wave Runup,"

Journal of Waterway, Port, Coastal, and Ocean Engineering, 124(4): 157- 171, 1998.

2.6-43 Not Used.

2.6-44 "Regional Assessment of Tsunami Potential in the Gulf of Mexico," U.S. Geological Survey Administrative Report, September 2, 2009.

2.6-45 "Global Centroid-Moment-Tensor Project." Available at http://www.globalcmt.org/,

accessed November 6, 2012.

2.6-46 "National Geophysical Data Center/World Data System (NGDC/WDS) Global Historical Tsunami Database," Boulder, CO, USA. Available at http://www.ngdc.noaa.gov/hazard/tsudb.shtml, accessed November 6, 2012.

2.6-47 STPEGS (Units 1 & 2) Updated Final Safety Analysis Report (UFSAR), Section 2.4, "Hydrologic Engineering," Rev. 15.

2.6-48 Chaytor, J. D., D. C. Twichell, P. Lynett, and E. L. Geist, "Distribution and Tsunamigenic Potential of Submarine Landslides in the Gulf of Mexico," in Submarine Mass Movements and Their Consequences, Advances in Natural and Technological Hazards Research, edited by D.C. Mosher et al., Springer Science +

Business Media B.V., 28: 2010.

2.6-49 South Texas Project Units 3 & 4 Combined License Application (COLA), Final Safety Analysis Report (FSAR), Rev. 7, Nuclear Innovation North America LLC (NINA),

February 1,2012.

2.6-50 Prasad, R, L. F. Hibler, A. M. Coleman and D. L. Ward, and Pacific Northwest National Laboratory, "Design-Basis Flood Estimation for Site Characterization at Nuclear Power Plants in the United States of America," NUREG/CR-7046, PNNL-20091, Nuclear Regulatory Commission, November 2011.

2.6-51 National Geophysical Data Center (NGDC), "Volcano Location Database Search,"

National Oceanic and Atmospheric Administration (NOAA). Available at http://www.ngdc.noaa.gov/nndc/struts/form?t=102557&s=5&d=5, accessed May 2, 2011.

Tsunami 2.6-20

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushlima Response Project 2.6-52 Bourgeois J., T. A. Hansen, P. L. Wiberg, and E. G. Kauffman, "ATsunami Deposit at the Cretaceous-Tertiary Boundary in Texas," Science, 241: 567-570, July 29, 1988.

2.6-53 Gale, A. S., "The Cretaceous-Palaeogene Boundary on the Brazos River, Falls County, Texas: Is There Evidence for Impact-Induced Tsunami Sedimentation?"

Proceedings of the Geologists Association, 117: 173-185, 2006.

2.6-54 Schulte, P., R. Speijer, M. Hartmut, and A. Kontny, "The Cretaceous-Paleogene (K-P) boundary at Brazos, Texas: Sequence Stratigraphy, Depositional Events and the Chicxulub Impact," Sedimentary Geology 184: 77-109, 2006.

Tsunami 2.6-21

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 FukushiniaResponse Project Table 2.6-1 Areas of Potential Seismic Tsunamigenesis in the Caribbean (Reference 2.6-3, pp. 105 and 107)

Caribbean Source Latitude (ON) Longitude (OW)

North Panama Deformation Belt 9-12 83-77 Northern South American 11.5-14 77-64 Convergent Zone Table 2.6-2 Source Parameters for Veracruz Scenario Rupture Length Width Depth Strike Dip Rake Max slip Epicenter Mw (km) (kin) (km) (degree) (degree) (degree) (m) 200 N,{__

200 8.2 200 70 5

135 20 2650 E 90 2 Table 2.6-3 Initial Wave Deformation Characteristics and Maximum Runup for Simulations Deformation Dipole Initial Dipole Initial Maximum Area (sq. km) Minimum (m Maximum (m Runup (m Figure Case below MSL) below MSL) above MSL) Reference PV 411 -7 3 1 2.6-7; 2.6-8 PV(x20) 387 -140 60 2 2.6-9; 2.6-10 PNG 879 -20 16 2 2.6-11 2.6-12 Monster 9932 -38 27 2 2.6-13; 2.6-14 Tsunami 2. 6-22

Enclosure NOC-AE-13002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Figure 2.6-1 Location Map of STP 1 & 2 from the Gulf Coast and Colorado River Tsunami 2. 6-23

Enclosure NOC-AE-1 3002975 flooding Hazard Reevaluation Report STPJI & 2 Fukushima Response Project EXPLAN4ATION A PrOPOSed USGS GSN Stabaft A Exrist USGS GSN Estdons

Proposed NOAA DART Stkmos

- PIMt boujndares 4, Volcanoes Earaquake spicenterv 1010-2004. Ma 6

Q-69 In 0 70-299

ao~e ev 1530-1991 Figure 2.6-2 Regional Map of Plate Boundaries and Tsunami-Generating Earthquakes from 1530-1991 in the Caribbean Sea (modified from Reference 2.6-21).

Tsunami 2.6-24 262

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project I

- - - ,oo~ ~

Note: At 142 km from STP I & 2, the East Breaks slump is the only nearfield landslide source. Source of bathymetry: Reference 2.6-41.

Figure 2.6-3 Landslide Source Regions in Gulf of Mexico Tsunami 2. 6-25 I

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project OM

-20 .

-LO00O, SLM k%IIINI V.ku ('(#N10 :

-1,500.

Note: Buoy record for recording tsunami wave amplitudes is located at 28.580 N and 95.98' W. Bathymetry elevations are relative to Mean Sea Level (MSL).

Figure 2.6-4 Location of East Breaks Slump Relative to STP 1 & 2 (Source: Reference 2.6-41)

Tsunami 2. 6-26

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project Note: Bathymetry elevations are relative to MSL. (Source of bathymetry data: Reference 2.6-41)

Figure 2.6-5 Source Parameters for East Breaks Slump Tsunami 2.6-27

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project I II Note: Bathymetry elevations are relative to MSL (Source of bathymetry data: Reference 2.6-41).

Figure 2.6-6 Grid Spacing for East Breaks Slump Modeling with MOST Tsunami 2.6-28 E

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Initial Condition - Palos Verdes (PV) Meters 27.9 0

.200

~27 .400

-0

-600

.800

.127,

.1000

-1200

-1400 27.

96 95.5 95 Longitude (degrees west)

Note: Elevations of initial wave correspond with elevations in Figure 2.6-8.

Figure 2.6-7 Plan View of Palos Verdes (PV) Initial Deformation Condition at Location of the East Breaks Slump in the Gulf of Mexico Tsunami 2. 6-29

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 4 ........................................................................................

2 2 ...................... ........... .......... ........... ........... ........ ...... ......

1 0 . ... ... .

4C it

-2

-4

..... : ........... .................... ..... t......... .

.i 264 . " .. ...... .......i ............ ::

2I II I I I I 2 27.7 27.68 27.66 27.64 27,62 27.6 27.58 27.56 27.54 27. 52 LAT1TIDE (DEG N)

LONGITUDE (DEG)

Note: Maximum elevation of negative wave is -7 m (MSL); maximum elevation of positive wave is +3 m. (MSL).

Figure 2.6-8 Side View of Palos Verdes (PV) Initial Deformation Condition Tsunami 2.6-30

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project initisii reaucihn - Pskine VrrAc v9O tPVv9Al 27.9, MectCrs 0

200

~27. .400

-600

- 800 327.

-1000

-1200

-1400 27.0!l 96.5 96 95.5 95 Longitude (degrees west)

Note: Elevations of initial wave correspond with elevations in Figure 2.6-10.

Figure 2.6-9 Plan View of Palos Verdes x20 (PVx20) Initial Deformation Condition at Location of the East Breaks Slump in the Gulf of Mexico Tsunami 2. 6-31

Enclosure NOC-AE-1 3002975 STP 1 & 2 Flooding HazardReevaluation Report Fukushima Response Project 60 20 20 ...................

..................... . ,400 120 0

S-20

: "'.:- 20

. ............................................... i .

M 0 N.. ..

W

-40 l........... ..... . .. .

80

-80 .... .. ......... ........... ..........

10 0

.... '" "00

-1 0............ ...........

"l00 °. . .......... ... ...... .............

-140 ... ...

""""""."... ........ ...... ................... :................ :"".... :.12 2 6 4 .4'. . ....

.... ... ... I . ... .... ...

....... ..... 120 264.35" " """. . ... .* : ' . ."

243 27.7 27.65 27.6 27.55 21 LONGITUDE (DEG) LATITUDE (DEG N)

Note: Maximum elevation of negative wave is -140 m (MSL): maximum elevation of positive wave is +60 m (MSL).

Figure 2.6-10 Oblique View of Palos Verdes x20 (PVx20) Initial Deformation Condition Tsunami 2.6-32

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Initial Condition - Papua New Guinea (PNG) Meters 27.8.

0

-200

-400 02 U

U

-600 U

3 -800

-1000

-1200

-1400 95.6 93.4 I, Longitude (degrees west)

Note: Elevations of initial wave correspond with elevations in Figure 2.6-12.

Figure 2.6-11 Plan View of Papua New Guinea (PNG) Initial Deformation Condition at Location of the East Breaks Slump in the Gulf of Mexico Tsunami 2.6-33

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project 20\

15 15 10 I* 5 0:. .... ...... ..... ....

M 0 3:

-5

.5

-15 .... ......... -7 ................................... ............

-10

-1

-20 .......

............ -15

'f".°............................. ...........

264.4 ...........

264.2 --

27.7 27.68 27.6 6 27.64 27.62 27.6 27.50 27.56 27.54 27.52 27.5 LATITLDE (DEG N)

Note: Maximum elevation of negative wave is -18 m (MSL); maximum elevation of positive wave is +16 m (MSL).

Figure 2.6-12 Oblique View of Papua New Guinea (PNG) Initial Deformation Condition Tsunami 2.6-34 I

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project initisil (nntfifinn - Mnntt-,r Meters 27.9 0

200

~27. 400 600 800

~27.

1000 1200 1400 27.

Longitude (degrees west)

Note: Elevations of initial wave correspond with elevations in Figure 2.6-14.

Figure 2.6-13 Plan View of Hypothetical "Monster" Initial Deformation Condition at Location of the East Breaks Slump in the Gulf of Mexico Tsunami 2. 6-35

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project 30-

. 20 20, 10 I0

: 0 I

S .......

-10

-20,

...... ..... -20

-30s ......r..-:.-'. .,°-.

40, 27.4 27... 6: ".-::."'".;.-. :: ......

-30 264'.8 '" 264.6 244 6. 6 27.8 .. "

LATITtXE (DEG 1J LONGITUDE (DEG)

Note: Maximum elevation of negative wave is -38 m (MSL); maximum elevation of positive wave is +28 m (MSL).

Figure 2.6-14 Oblique View of Hypothetical "Monster" Initial Deformation Condition Tsunami 2.6-36 I

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project 00

.,500,

-1,o0WO

-1,250=m I

10 m Figure 2.6-15 Maximum Coastal Runup for the PV Simulation Tsunami 2.6-37

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project BUOY RECORD AT 8. 1M DEPTH 11 1 0.2 Z

0o.1

-~0 Li-L-0.

-0.

-0.3 0.5 1 1.5 2 2.5 3 3.5 TIME (HOURS)

Note: Datum referenced to MSL.

Figure 2.6-16 Time Series of Wave Amplitude for PV Simulation at 28.580 N and 95.980 W (i.e., Buoy Location Shown in Figure 2.6-4)

Tsunami 2. 6-38 I

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project OM

-"Dm --

-50, -

-1,000M l.0O - II Figure 2.6-17 Maximum Coastal Runup for the PVx20 Simulation Tsunami 2.6-39

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project BUOY RECORD AT 8.1 M DEPTH H

U 0.5 1 1.5 2 2.5 3 3.5 TIME (HOURS)

Note: Datum referenced to MSL.

Figure 2.6-18 Time Series of Wave Amplitude for PVx20 Simulation at 28.58* N and 95.98"W (i.e., Buoy Location Shown in Figure 2.6-4)

Tsunami 2.6-40

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project Ow

-250 =

-5003 1500, 10im Figure 2.6-19 Maximum Coastal Runup for the PNG Simulation Tsunami 2.6-41

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project BUOY RECORD AT 8.1 M DEPTH

- 0.8 Z 0.6 H

< 0.4 0.2 Ud

< 0

-0.2 LL-0.

-0.6 0.5 1 1.5 2 2.5 3 3.5 TIME (HOURS)

Note: Datum referenced to MSL.

Figure 2.6-20 Time Series of Wave Amplitude for PNG Simulation at 28.580 N and 95.98* W (i.e., Buoy Location Shown in Figure 2.6-4)

Tsunami 2. 6-42

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP & 2 Fukushima Response Project Ou. STP1 2

-500m -

.h.

-1I00WM-

-1250m --

.l*6m ..A l

-1 lm l0bn Figure 2.6-21 Maximum Coastal Runup for the Hypothetical "Monster" Simulation Tsunami 2.6-43

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project BUOY RECORD AT 8.1 M DEPTH 0O.5

.- 0 2

-2.5 0.5 1 1.5 2 2.5 3 3.5 TIME (HOURS)

Note: Datum referenced to MSL.

Figure 2.6-22 Time Series of Wave Amplitude for Hypothetical "Monster" Simulation at 28.680 N and 95.980 W (i.e., Buoy Location Shown in Figure 2.6-4)

Tsunami 2. 6-44 I

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushinza Response Project 2.7 Ice Induced Flooding 2.7.1 Ice Conditions and Historical Ice Formation The potential impact of ice effects on the STP 1 & 2 site was analyzed by evaluating air and water temperature data in the vicinity of the site. A review of historical air temperature data indicates that the climate in the vicinity of the site is temperate. There is also no record of ice formation that could affect the safety-related facilities at the site.

Water temperature data recorded at three Lower Colorado River Authority (LCRA) stations on the Lower Colorado River were analyzed to determine minimum water temperatures in the river.

These LCRA stations include Bay City (Site #12284), Wharton (Site #12286), and Columbus (Site #12290) and they are located approximately 14, 37, and 71 miles from the STP 1 & 2 site, respectively (Reference 2.7-1). This data covers the period from 1982 through 2006 for the river reach below Mansfield Dam, which is located upstream of the STP site, approximately 156 miles northwest of the site. The recorded surface water temperatures at the selected LCRA stations show that the water temperature has remained above the freezing point during this period as shown on Figure 2.7-1. The minimum recorded daily water temperature at these stations was 41.2 0 F (5.1 0 C), which occurred on February 6, 1985. In addition, the UFSAR for STP 1 & 2 indicates that for water temperature data available between October 1944 and Septmenber 1975, a minimum water temperature of 35 OF occurred in December 1963 and January 1964 (Reference 2.7-7).

Subsequent review of water temperature data up to 2012 also showed that minimum water temperatures remained above freezing (Reference 2.7-2).

Long-term air temperature records available at the STP site and the Bay City climate station show that the air temperature at the plant site rarely drops below the water freezing point. When freezing temperatures do occur they do not persist for long periods of time. Daily data from the STP site for the period 1990 to 2006 show that there was only one instance where the daily average air temperature was below the water freezing point for three consecutive days (see Table 2.7-1). Daily data from the Bay City climate station for the period 1942 to 2006 (Reference 2.7-5) show that over this 65-year period, there was only one instance where the daily average air temperature was below the water freezing point for five consecutive days, two instances that the daily average air temperature was below the water freezing point for four consecutive days, and four instances that it was below the water freezing point for three consecutive days (see Table 2.7-2). These data suggest that conditions conducive to ice formation rarely occur, and that when they do, they do not persist for more than a few days.

Subsequent review of Bay City air temperature data through July 2012 (Reference 2.7-6) did not uncover any cold weather events that present a risk of ice formation with potential to impact the safety-related operations of the plant.

2.7.2 Ice Jam Events A search of the "Ice Jam Database" maintained by the U.S. Army Corps of Engineers (USACE) for records up to 2006 revealed no records of ice jams on the Lower Colorado River (Reference 2.7-3). A subsequent search of the database in 2012 also did not reveal any recorded ice jam events from 2006 to 2012 (Reference 2.7-4).

Ice-Induced Flooding 2. 7-1

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project 2.7.3 Conclusion As established in the subsections above, the results of the evaluation of historical air and water temperature data in the vicinity of STP Units 1 & 2 and search of the USACE Ice Jam Database indicate that ice induced flooding is not a credible hazard that will adversely impact the safety functions of the plant.

2.7.4 References 2.7-1 Lower Colorado River Authority (LCRA). Available at http://waterquality.lcra.org/,

accessed February 15, 2007.

2.7-2 Lower Colorado River Authority (LCRA). Available at http://waterquality.Icra.org/,

accessed August 1, 2012.

2.7-3 "Ice Jam Database," U.S. Army Corps of Engineers, Cold Region Research and Engineering Laboratory (CRREL). Available at http://www.crrel.usace.army.mil/ierd/ijdb/, accessed February 10, 2007.

2.7-4 "Ice Jam Database," U.S. Army Corps of Engineers, Cold Region Research and Engineering Laboratory (CRREL). Available at https://rsgisias.crrel.usace.army.mil/apex/f?p=273:2:85695751952301, accessed August 1, 2012.

2.7-5 NOAA National Climatic Data Center (NCDC). Available at http://cdo.ncdc.noaa.gov/CDO/dataproduct, accessed on January 31, 2007.

2.7-6 NOAA National Climatic Data Center (NCDC). Available at http://www.ncdc.noaa.gov/cdo-web/, accessed August 27, 2012.

2.7-7 STPEGS Updated Final Safety Analysis Report, Units 1 & 2, Section 2.4 (Hydrologic Engineering), Revision 15.

Ice-Induced Flooding 2.7-2

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Table 2.7-1 Lowest Average Daily Temperature and Number of Days with Average Daily Temperature below Freezing at STP Site Lowest Average Daily Date Lowest No. of Temperature Average Daily Consecutive Total No. of Temperature Freezing Freezing Year F °C Occurred Days Days 1990 27.2 -2.6 12/21190 2 2 1991 35.6 2.0 12/15/91 0 0 1992 42.1 5.6 7/11192 0 0 1993 35.9 2.2 10/25/93 0 0 1994 39.7 4.3 12/1/94 0 0 1995 37.3 3.0 11/8195 0 0 1996 26.4 -3.1 1/8/96 3 4 1997 30.9 -0.6 1/13/97 0 1 1998 35.5 1.9 12/25/98 0 0 1999 36.3 2.4 1/4/99 0 0 2000 36.6 2.5 12/12/00 0 0 2001 34.3 1.3 113/01 0 0 2002 35.5 1.9 1/2/02 0 0 2003 37.4 3.0 2/24/03 0 0 2004 32.6 0.4 12/24/04 0 0 2005 42.2 5.7 1/22/05 0 0 2006 38.6 3.7 2/18/06 0 0 Average (days) 0.3 0.4 Ice-Induced Flooding 2. 7-3

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushinia Response Project Table 2.7-2 Lowest Average Daily Temperature and Number of Days with Average Daily Temperature below Freezing at Bay City Climate Station Lowest Average Daily Date Lowest No. of Temperature Average Daily Consecutive Total No. of Temperature Freezing Freezing 0

Year OF C Occurred Days Days 1942 40.5 4.7 12/28/1942 0 0 1943 31.5 -0.3 1/2611943 0 1 1944 35.5 1.9 1/1411944 0 0 1945 35.5 1.9 12/20/1945 0 0 1946 34.5 1.4 12/30/1946 0 0 1947 28 -2.2 1/4/1947 1 2 1948 25.5 -3.6 1/29/1948 3 5 1949 25.5 -3.6 1/3011949 0 1 1950 28.5 -1.9 12/7/1950 0 1 1951 20 -6.7 2J211951 3 5 1952 42 5.6 11/30/1952 0 0 1953 30.5 -0.8 12/24/1953 0 1 1954 30 -1.1 1/22/1954 0 1 1955 36 2.2 3/27/1955 0 0 1956 34 1.1 2/4/1956 0 0 1957 33 0.6 1/17/1957 0 0 1958 30 -1.1 2/13/1958 0 1 1959 29.5 -1.4 1/511959 0 1 1960 33 0.6 2/2511960 0 0 1961 34.5 1.4 1/29/1961 0 0 1962 22.5 -5.3 1/11/1962 1 2 1963 24.5 -4.2 1/13/1963 3 7 1964 29.5 -1.4 1/14/1964 0 1 1965 34.5 1.4 2/25/1965 0 0 1966 28.5 -1.9 1/30/1966 0 1 1967 33.5 0.8 217/1967 0 0 1968 34 1.1 1/8/1968 0 0 1969 36 2.2 1/5/1969 0 0 1970 31 -0.6 1/711970 0 1 ice-hiduced Flooding 2. 7-4

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushimia Response Project Table 2.7-2 Lowest Average Daily Temperature and Number of Days with Average Daily Temperature below Freezing at Bay City Climate Station (continued) 1971 31.5 -0.3 1/8/1971 0 1 1972 32 0.0 1/5/1972 0 0 1973 25.5 -3.6 1/12/1973 4 5 1974 27 -2.8 1/4/1974 0 1 1975 27.5 -2.5 1/13/1975 0 1 1976 29.5 -1.4 11/29/1976 0 1 1977 31.5 -0.3 1/19/1977 0 1 1978 28 -2.2 1/21/1978 1 2 1979 26 -3.3 1/2/1979 1 3 1980 31 -0.6 3/2/1980 0 1 1981 30.5 -0.8 2/12/1981 0 1 1982 27 -2.8 111411982 0 3 1983 20.5 -6.4 12/25/1983 5 6 1984 31 -0.6 1/20/1984 1 2 1985 23.5 -4.7 2/2/1985 3 6 1986 36.5 2.5 2/12/1986 0 0 1987 56.5 13.6 2110/1987 0 0 1988 42.5 5.8 12/17/1988 0 0 1989 16.5 -8.6 12/23/1989 4 6 1990 23 -5.0 12/23/1990 1 2 1991 35 1.7 1/1/1991 0 0 1992 36.5 2.5 11/27/1992 0 0 1993 34 1.1 11127/1993 0 0 1994 34 1.1 212/1994 0 0 1995 40 4.4 1/5/1995 0 0 1996 31.5 -0.3 1/8/1996 0 1 1997 31.5 -0.3 1/14(1997 0 1 1998 35.5 1.9 12/24/1998 0 0 1999 36 2.2 1/3/1999 0 0 2000 37.5 3.1 12/13/2000 0 0 2001 34.5 1.4 1/2/2001 0 0 2002 35 1.7 2127/2002 0 0 Ice-Induced Flooding 2. 7-5

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STPI & 2 Fukushinma Response Project Table 2.7-2 Lowest Average Daily Temperature and Number of Days with Average Daily Temperature below Freezing at Bay City Climate Station (continued) 2003 40.5 4.7 1/18/2003 0 0 2004 33.5 0.8 12/26/2004 0 0 2005 33 0.6 12/9/2005 0 0 2006 40.5 4.7 2119/2006 0 0 Average (days) 0.5 1.2 Ice-Induced Flooding 2.7-6

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP 1& 2 Fukushima Response Project Water Temperatures in Lower Colorado River 100, 90

i I t ---

80 90A* ,i* *I.* : g I' ----* " '" 1 aAt Wharton 50 ----- .- *-;i l-*-' - " -, --- .---- ..- "-------t

In(I

,,, , -,-- i- ...* - . . ,

I-------- *. . .. o At Columbus

At BayCity if 60 a I a 0 I

"- - = - ,' Freezing temperature (32 0 F) ,---

3 0 -- - -- - -- - - - . . . . . . . . . . -- .. -

Minimum recorded 10------------------------- temperature =-41. -- ----------------

0 02,/18/82 11/14/84 08/1'1/87 05/0'7/90 01/31/93 10/2'8/95 07/2-4/98 04/1'9/01 01/1'4/04 10/1'0/06 Figure 2.7-1 Recorded Water Temperatures in the Lower Colorado River Ice-Induced Flooding 2.7-7

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project 2.8 Channel Migration or Diversion 2.8.1 Historical Channel Migration or Diversions STP 1 & 2 is located adjacent to the Lower Colorado River approximately 16 miles upstream from the Gulf of Mexico in Matagorda County, Texas (Figure 2.8-1). The geologic history of Matagorda County within the last 200,000 years has been studied extensively by different investigators (References 2.8-1, 2.8-2, and 2.8-3). The oldest and most prominent geologic formation in Matagorda County is the Beaumont Formation (Figure 2.8-2). The Beaumont Formation is described in Reference 2.8-1 (p. 112-113) as "a regressive or prograding sedimentary geologic unit" that was "probably laid down as an alluvial plain by a paleo-Colorado River. The ancient river's successive meandering courses distributed fluvial and deltaic sediments between the contemporaneous Pleistocene alluvial plains of the paleo-Lavaca and paleo-Navidad Rivers to the southwest and a paleo-Brazos River to the northeast." Early delineations of the Lavaca, Navidad, Colorado, and Brazos rivers are shown in Figure 2.8-3, which is a reproduction of an 1838 map of Texas showing the Rio Colorado, i.e., the Lower Colorado River, and other nearby rivers (Reference 2.8-4). The historical Colorado channel followed the same course as the present day Caney Creek channel, which is referred to as Cane Brake in the map.

Reference 2.8-1 (p. 112) states that the Beaumont Formation was "deposited during a late Pleistocene high sea-level stand similar to that of the present." However, the interglacial period during which the deposition took place is debated. Reference 2.8-1 (p. 112) states that "some investigators place its deposition during the Sangamon [i.e., between 110,000 and 130,000 years ago], a major interglacial stage between the Illinoian [i.e., between 130,000 and 200,000 years ago] and Wisconsin glacial stages [i.e., about 12,000 to 110,000 years ago]. It is possible that the formation was deposited less than 35,000 years BP (before present) in a late intra-Wisconsin high sea-level stand." With respect to the mechanism for deposition, Reference 2.8-1 (p. 112) states that "during the Pleistocene, when the continental glaciers expanded several times, water was transferred from the ocean basins to the largely land-based glaciers and there was a worldwide lowering of sea level. Estimated sea levels below present-day levels range from about 250 to 450 feet. Streams draining into the oceans incised and regraded their channels as they flowed toward a lower, more distant seashore. When the sea level rose in response to periodic melting of the glaciers, the incised channels were flooded, backfilled or alluviated. Subsequently, broad alluvial plains were built along the gulf coast."

The Quaternary geologic history of the paleo-Colorado River has been less well documented upstream of Matagorda County than in Matagorda County. Reference 2.8-2 (p. 1003) states that the Colorado River is "draining the geologically heterogeneous Southern High Plains and Edwards Plateau regions of West Texas. As the channel emerges from a deep canyon at the Balcones Escarpment, the drainage basin narrows considerably, and the lower Colorado River transects the Gulf Coastal Plain for 280 km until discharging into the Gulf of Mexico. On the inner coastal plain, the lower Colorado River flows within a well-defined bedrock valley that transects Upper Cretaceous carbonates, then progressively younger and less steeply dipping Tertiary siliciclastic rocks."

From the late Quaternary (i.e., between approximately 200,000 years to 35,000 years ago) to the recent present (i.e., approximately 550 years ago), the Colorado River channel occupied the present day Caney Creek channel (Figure 2.8-1, References 2.8-1 and 2.8-2). As stated in Reference 2.8-1 (p. 11), "the flood plain along an abandoned Holocene course of the Colorado Channel Migration or Diversion 2.8-1

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushimna Response Project River, now occupied by Caney Creek, lies along the northeastern margin of the county."

Stratigraphy investigations indicate a recent major avulsion (i.e., the shifting of flow from one channel to another channel) of the Colorado River occurred near Glen Flora, Texas, approximately 500 to 600 years ago (Reference 2.8-2). The previous channel belt (i.e., the previous river channel occupied by the Colorado River before the river avulsed) flowed southeast down the modern day Caney Creek channel and into the eastern end of Matagorda Bay (Figure 2.8-1). The two channels are no longer connected (Figures 2.8-2 and 2.8-3). A delineation of the historical Colorado channel is presented in Figure 2.8-1.

The Caney Creek meander belt and geomorphic features are relatively mature compared to the present day Colorado River channel. Reference 2.8-2 (p. 1009-1010) states that "the Caney Creek meander belt was substantially different from the modern channel and suggested that it was a highly sinuous, mature, fully aggraded channel course prior to its abandonment (Figure 2.8-1). Indeed, examination of the Caney Creek meander belt in air photos and in the field shows that a well-defined levee, crevasse splay, and flood basin depositional environments were common to the lower Colorado River when itflowed through the Caney Creek course.

Such features did not occur along the lower Colorado River in the bedrock-confined portion of the valley, upstream from the point of avulsion, or in the recently occupied channel farther downstream until the lowermost reaches near the present shoreline." Further, Reference 2.8-2 (p. 1014-1015) states that "avulsion and abandonment of the fully aggraded Caney Creek meander belt, with occupation of the modern lower Colorado channel, most likely occurred in response to near complete filling of the incised valley during the present highstand."

2.8.2 Stratigraphic Evidence Stratigraphic formations along the Colorado River and Caney Creek indicate the river has occupied the recent course near STP Units 1 & 2 for approximately the last 550 years.

Stratigraphic records also indicate that, in an unregulated setting, the most likely zone for future river avulsion is between Eagle Lake and Wharton (Figure 2.8-1). Reference 2.8-2 (p. 1009) states "between Eagle Lake and Wharton, the abandoned Caney Creek meander belt and modern channel occur within a single valley that contains the ELA (i.e., the Eagle Lake Alloformation, which was deposited approximately 20,000 to 14,000 years BP), CBA-1 [i.e., the Columbus Bend Alloformation, which was deposited approximately 12,000 to 5,000 years BP],

CBA-2 [i.e., the Columbus Bend Alloformation, which was deposited approximately 5,000 to 1,000 years BP] and floodplain facies from the Caney Creek meander belt and the modern channel" (Figure 2.8-4). Further, "downstream from Wharton, the Caney Creek and modern channel courses diverge and ultimately discharge into the Gulf of Mexico some 40 km from each other" (Figure 2.8-1). Reference 2.8-2 (p. 1015) states "the influence of base-level change on stratigraphic architecture in the lower Colorado valley extended 90 km upstream from the present highstand shoreline [near Eagle Lake, Texas], but was superimposed on climatically driven episodes of sediment storage or removal. Thus, depositional sequences on coastal plain rivers with large inland drainage basins most likely record interactions between upstream controls on discharge and sediment supply, and base-level controls on stratigraphic architecture and preservation in the geologic record, rather than a strict one-to-one relationship with base-level change per se."

Channel Migration or Diversion 2.8-2

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fk~ushima Response Project 2.8.3 Ice Causes With respect to ice, there is no record of any major river in the State of Texas freezing over at any time in recorded history (Section 2.7). Consequently, ice jams that could cause a channel migration or diversion are considered unlikely.

2.8.4 Flooding of Site Due to Channel Migration or Diversion 2.8.4.1 Geologic Effects A record of channel diversions due to upstream and above-bank channel changes due to geologic, seismic, or topographic changes, including subaerial landslides and earthquakes, has not been documented above the transition into the Balcones Escarpment, which occurs near Austin, Texas (Reference 2.8-1). As stated on p. 1003 of Reference 2.8-1, "the Colorado River is a large fluvial system [...] with its upper reaches and all major tributaries (92% of total area) draining the geologically heterogeneous Southern High Plains and Edwards Plateau regions of West Texas. As the channel emerges from a deep canyon at the Balcones Escarpment, the drainage basin narrows considerably, and the lower Colorado River transects the Gulf Coastal Plain for 280 km until discharging in the Gulf of Mexico." In the vicinity of STP Units 1 & 2, the region is relatively flat, with less than a one degree average dip in regional geologic units from the location of STP Units 1 & 2 to the Gulf of Mexico (Reference 2.8-5). This low dip indicates a low probability of slope failure along bedding planes. While growth faults are common geologic structures in the Texas Coastal Plain, these faults are non-tectonic gravity-related displacements formed within sediment deposition of the geologic formations. The information presented in References 2.8-6a through 2.8-6f indicate that there are no capable faults in the STP site region. Therefore, it is highly unlikely that surface faulting can occur and cause a slope failure that would lead to channel diversion or surface faulting that would displace landforms and, thereby, cause channel diversion.

2.8.4.2 Land Subsidence from Groundwater Pumping In the vicinity of Bay City, the measured land subsidence for the period of 1918 to 1951 was only 0.12 ft (Reference 2.8-7a). Most of the 1918 to 1951 subsidence may be attributed to increased use of groundwater after 1940 (Reference 2.8-7a). From 1943 to 1973, the land surface subsided more than 1.5 ft (0.46 m) due to groundwater withdrawals (Reference 2.8-7b).

More recently, however, land subsidence in Matagorda County has been relatively minimal due to declining groundwater use. Reference 2.8-7c (p. 7) states the groundwater use for Matagorda County as 38,554 acre-feet in 1980, 37,537 acre-feet in 1990, and 14,413 acre-feet in 1997. In 1997, less than 10% of total water use was derived from groundwater sources (Reference 2.8-7c, p. 7).

2.8.4.3 Floods Of the various mechanisms that could cause channel migration or diversion, the most likely scenario for a major channel avulsion would be from a large flood, a series of large floods, the failure of upstream dams, or significant sea-level change. In an unregulated setting, the most likely location for a channel migration or diversion on the Colorado River would be between Eagle Lake, Texas, and Wharton, Texas (Figure 2.8-1). However, flows on the Lower Colorado River have been regulated since 1938. For example, since the completion of Lake Buchanan (1937) and Lake Travis (1940), the peak discharge for the Colorado River at Austin (USGS Gauge #08158000) was 47,600 cubic feet per second (cfs) in April 1941 (Figure 2.8-5). A flood Channel Migration or Diversion 2.8-3

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project that occurred in September 1952 would have produced a flow of over 700,000 cfs had Mansfield dam and Lake Travis not been present. However, Lake Travis has sufficient storage capacity to withhold the entire flood volume. Instead of a potentially disastrous flood, the peak discharge recorded at Austin during this period was only 3720 cfs (Reference 2.8-6a).

2.8.4.4 Erosion and Channel Diversion due to Coastal Storm Surges The largest documented hurricane to impact the Texas Coast was Hurricane Carla in 1961 (Section 2.4). Reference 2.8-1 (p. 114) states that "Hurricane Carla partly obliterated Matagorda Peninsula in 1961. Erosion effects, however, were soon repaired by shoreline deposition and wind-driven migration of shoreline sediments across the peninsula." Further, "many scoured washover or storm channels eroded during hurricanes are transverse to the general trend of the peninsula. Almost all are sealed from the gulf by the present-day beach." Because Hurricane Carla was nearly equivalent to the Probable Maximum Hurricane discussed in Section 2.4, hurricane effects are not considered to be a significant mechanism for channel diversion that would impact the safety function at STP Units 1 &2.

2.8.4.5 Channel Diversion to Upstream Gravel Mining Effects Sand and gravel mining activities in the Colorado River occur in the vicinity of and immediately downstream of Austin, Texas (Reference 2.8-8). Reference 2.8-8 (p. 883) states that "flooding has caused the river to erode its banks and carve new paths through abandoned pits, effectively altering the river course at several locations in Travis and Colorado counties in Texas."

Reference 2.8-8 notes that gravel mining has led to artificial cutoffs of historical river meanders and localized downstream bank effects. In addition, Reference 2.8-9 notes that "gravel mining without appropriate constraints can lead to severe bed degradation downstream, with the resulting failure of bridges [and] exposure of buried pipelines." However, severe bed degradation effects in the Lower Colorado River have not been documented. Consequently, gravel mining effects are not considered to be a significant mechanism for channel diversion that would impact the safety function of STP Units 1 & 2.

2.8.5 Human-Induced Changes of Channel Diversion 2.8.5.1 Colorado River Delta The geomorphology of the Lower Colorado River since the late 17th century was largely governed by the occurrence and subsequent removal of a major log jam blocking the river near Wharton, Texas. Reference 2.8-3 (p. 100) states that "the earliest historical reference to a raft of logs in the Colorado River was made in 1690 when the Matagorda Bay area was mapped by Spanish explorers headed by Captain Francisco de Llanos. Deposition of the modern delta must have begun after 1690 because the Spaniards were able to ascend some 10 or 15 miles of the eastern channel of the Colorado River, which is about at the head of the tidewater. Had there been deltation at the mouth of the river, it is unlikely that they could have gotten their sailing ships into the river. The Spaniards discovered a log raft (i.e., debris jam) in the western channel and had to turn back and exit by another channel. This raft was mentioned by William Selkirk in 1824, the first surveyor in the area." Reference 2.8-1 (p. 115) states that "in 1824, the downstream edge of the raft was about 46 miles in length and entered Wharton County.

Unsuccessful and poorly funded efforts to destroy the raft persisted until 1925, when a narrow pilot channel was blasted through the raft. A major flood on the Colorado River in 1929 carried substantial parts of the raft into Matagorda Bay and silted up the mouth of the river channel. The Channel Migration or Diversion 2.8-4

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project municipality of Matagorda and the surrounding lowlands were then subjected to periodic flooding. During the flood of 1935, the major flow of the Colorado River was almost diverted into Tres Palacios Creek and Tres Palacios Bay, one of the arms of Matagorda Bay." Reference 2.8-3 (p. 103) states "the last major flood occurred in 1935, when considerable water from the Colorado River found its way into the head waters of Tres Palacios Creek in Wharton County. If left alone, the Colorado River would have diverted itself again and Tres Palacios Creek might be now the main channel of the Colorado River." Further, "concurrent dam building and flood control measures in the upper Colorado watershed greatly reduced the danger of flooding in the Colorado lowlands."

The removal of the log raft led to the development of the Colorado River delta near Matagorda (Figure 2.8-6). The development of the delta eventually separated East Matagorda Bay from Matagorda Bay. In 1908, the delta spanned 45 acres. In 1929, a large flood flushed much of the raft into Matagorda Bay, and silted up the mouth of the channel, splitting the bay into East Matagorda Bay and Matagorda Bay. The acreage of the delta increased to 3470 acres in 1933, 4890 acres in 1936, 7098 acres in 1941, and 7200 acres in 1953 (Reference 2.8-3, Figure 2.8-7). In 1936, a channel was cut through the peninsula to relieve flooding, and the Colorado River discharged directly into the Gulf of Mexico. Since 1941, Mansfield Dam and Buchanan Dam have trapped most of the coarse sediment in the Colorado River. Consequently, the delta has been in a recessive mode (Reference 2.8-1).

2.8.5.2 Channel Stabilization and Efficiency Constructed channels often define channel efficiency in terms of the channel cross-section that "gives the maximum discharge, Q, for a specified flow area, A," which is known as the most efficient hydraulic section (Reference 2.8-10, p. 235). However, in natural channels, channel efficiency refers to the effective ability of the channel to move both water and sediment over a wide range of flows and grain sizes, respectively. Reference 2.8-11 (p. 168) states that "rivers with erodible boundaries flow in self-formed channels that, when subject to relatively uniform controlling conditions, are expected to show a consistency of form, or average geometry, adjusted to transmit the imposed water and sediment discharges." The adjustment of the local channel form is a function of the shear stress relative to sediment supply.

With respect to sediment transport, erosion, deposition and longitudinal profile impacts due to Lake Buchanan and Lake Travis, Reference 2.8-9 (p. 7) states that "the installation of a dam on a river typically blocks the downstream delivery of all but the finest sediment, creating a pattern of bed aggradation upstream. The dam raises base level, i.e., the downstream water surface elevation to which the river upstream must adjust, forcing upstream-migrating deposition."

Further, Reference 2.8-9 (p. 8) states that "the cutoff of sediment at a dam often induces bed degradation, as the river mines itself to replace the lost load. Bed degradation rarely continues unabated. Even small amounts of coarse, erosion resistant material in the substrate tend to concentrate on the bed surface as the bed degrades, eventually limiting the process through the formation of a static armor."

A considerable number of channel efficiency improvements were completed by the United States Army Corps of Engineers (USACE) south of Bay City in connection with the navigation project authorized by Congress under Section 7 of the Rivers and Harbors Act of August 8, 1917. Dredging was carried out between river mile 22 and the Gulf Intracoastal Waterway. This dredging stabilized the river planform (i.e., the lateral footprint of the channel) (Reference 2.8-6a). The dredged material was deposited along both banks of the river and the spoil areas were Channel Migration or Diversion 2.8-5

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project enclosed by embankments, limiting alluvial channel meander changes. A considerable portion of the abandoned river channel north of the STP Units 1 & 2 and in the vicinity of Selkirk Island was filled in shortly after 1917. Hence, shifting of the Colorado River channel near the project site is unlikely (Reference 2.8-6a).

In the vicinity of the site, natural levees have been developed along both banks of the Colorado River. Near the highway bridge FM 521 bridge crossing at river mile 16, the elevation of the levee is approximately 20 ft. Based on historical data collected by the USACE, a flood of 75,000 cfs would overtop the west bank near the site for existing channel and flood plain conditions.

The natural levee reaches an elevation of 25 ft approximately 2 miles upstream from FM 521. At this point the discharge required to overtop the levee under previous conditions was also approximately 75,000 cfs. Past backwater studies indicate that the bankfull capacity in the vicinity of STP Units 1 & 2 has increased to approximately 100,000 cfs, due in part to the dredging of a 14-foot-deep channel with a 100-ft width for a distance of 15.5 miles above the Gulf Intracoastal Waterway (Reference 2.8-6a).

Numerous relict Colorado River channels have also been documented from the Tres Palacios River west of the Colorado River to Caney Creek east of the Colorado River (Reference 2.8-1, Figure 2.8-1). For example, three miles downstream from Wharton, Texas, a west branch of the Colorado River formerly diverted flows south to Matagorda Bay (Figures 2.8-8 and 2.8-9).

Access to the west branch from the main river course was terminated when dredge spoil was used to fill in the connection in 1917 (Reference 2.8-6a). During flood stage, the west branch still conveys some of the overbank flows. The two branches isolate an island known as Wild Cow Island (Figure 2.8-8). Throughout Wild Cow Island, there are indications of abandoned river courses (Reference 2.8-1). To the east of the present Colorado River channel, Dick Island and Selkirk Island are formed by abandoned river courses (Figures 2.8-8, 2.8-9), some of which have also been blocked or filled by dredge spoil (Reference 2.8-6a).

2.8.5.3 Potential of Future Channel Migration and Impact The formation and evolution of an avulsion-dominated delta floodplain in which the Lower Colorado River flows is a complex process (Reference 2.8-12). Reference 2.8-13 (p. 711) states that "at present, little evidence is available on avulsion rates, avulsion frequencies, and inter-avulsion periods of aggrading fluvial systems over time scales of millennia." A river avulsion occurs with the rapid transfer of flow from the current channel to a new flow pathway. Studies of avulsion have noted several recurring characteristics, including persistent avulsion locations, the duration of inter-avulsion periods (i.e., the period of activity between channel belts), and avulsion frequency (Reference 2.8-13). In a study of another avulsion-dominated river system, persistent avulsion zones occurred in areas with a large difference in topographic elevation between the former flow course and the new flow course (Reference 2.8-13). Avulsion frequency has also been found to increase with increasing sedimentation rates that build or lead to aggradation of the channel relative to an adjacent flow course (Reference 2.8-14).

Consequently, before the construction of Lake Buchanan, Lake Travis and other upstream reservoirs, the potential for a channel diversion in an unregulated setting would be higher, especially if the sea level were to decline. However, upstream reservoirs have significantly attenuated large floods and trapped all but the finest sediment loads. In conjunction, the Lower Colorado River has had significant levees constructed along its length that has stabilized the river planform.

Channel Migration or Diversion 2Z8-6

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project The above review of the evidences and potential causes of channel diversions in the Colorado River indicates that there is little likelihood that major channel diversions impacting the safety facilities and function of STP Units 1 & 2 would occur. Specifically, flooding events in the order of a PMF at the STP site discussed in Section 2.2 as a result of channel diversions is considered improbable. Similarly, interruption of the non-safety water supply to the STP Reservoir Makeup Pumping Facility located on the west bank of the Colorado River as a result of channel migration is considered unlikely.

2.8.6 References 2.8-1 "Soil Survey of Matagorda County, Texas," Hyde, H. W. 2001, United States Department of Agriculture and Natural Resources Conservation Service.

2.8-2 "Late Quaternary Sedimentation, Lower Colorado River, Gulf Coastal Plain of Texas," Geological Society of America Bulletin 106: 1002-1016, Blum, M.D. and S.

Valastro, Jr., 1994.

2.8-3 "Historical Deltation of the Colorado River, Texas" Deltas in Their Geologic Framework, Houston Geological Society, p.99-105, Wadsworth, Jr., A. H., 1966.

2.8-4 "An Illustrated Atlas, Geographical, Statistical, and Historical, of the United States and the Adjacent Countries," Bradford, T. G., 1838.

2.8-5 "The Ouachita orogenic belt," The Geology of North America, v. F-2, The Appalachian-Ouachita orogen in the United States, pp.555-561, Viele, G. W.,

Geological Society of America, 1989.

2.8-6a STPEGS (Units 1 & 2) Updated Final Safety Analysis Report (UFSAR), Section 2.4, "Hydrologic Engineering," Rev. 15.

2.8-6b "Signatures of Climate vs. Sea-Level Change within Incised Valley-Fill Successions:

Quaternary Examples from the Texas Gulf Coast," Sedimentary Geology, v. 190, p.

177-211, Blum, M.D., and Asian, A., 2006.

2.8-6c "Seismic Hazard Methodology for the Central and Eastern United States,Tectonic Interpretations," v. 5 through 10, Electric Power Research Institute (EPRI), 1986.

2.8-6d "Data For Quaternary Faults, Liquefaction Features, and Possible Tectonic Features in The Central And Eastern United States, East of the Rocky Mountain Front,"

United States Geological Survey Open-File Report 00-260, Crone, A.J., and Wheeler, R.L., 2000.

2.8-6e "Known or Suggested Quaternary Tectonic Faulting, Central and Eastern United States-New and Updated Assessments for 2005," U.S. Geological Survey Open-File Report 2005-1336, Wheeler, R.L., 2005.

2.8-6f "Fault Number 924, Gulf-Margin Normal Faults, Texas," Quaternary fault and fold database of the United States, U.S. Geological Survey, Wheeler, R.L., compiler, 1999. Available at http://earthquakes.usgs.govfregional/qfaults, accessed May 16, 2007.

Channel Migration or Diversion 2.8-7

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 FukushitnaResponse Project 2.8-7a "Ground-Water Resources of Matagorda County, Texas," Report No. 91, Texas Water Development Board, March 1969.

2.8-7b "Land-Surface Subsidence in the Texas Coastal Region," Texas Water Development Board Report 272, Ratzlaff, Karl W., 1982.

2.8-7c "Aquifers of the Gulf Coast of Texas," Report 365, Texas Water Development Board, February 2006.

2.8-8 "Impacts of Sand and Gravel Mining on Physical Habitat of the Colorado River and Tributaries, Central Texas," Transactions of the Gulf Coast Association of Geological Societies, p. 883-890, Saunders, G. P., 2002.

2.8-9 "Transport of Gravel and Sediment Mixtures" ASCE Manual 110, Sediment Engineering, Chapter 3, Parker, G., 2007 (in press). Draft available at http://cee.uiuc.edu/people/parkerg/manual_54.htm, accessed June 21, 2007.

2.8-10 "Open-Channel Flow," Chaudhry, M. H., 1993.

2.8-11 "Fluvial Forms and Processes," Knighton, D., 1998.

2.8-12 "A Genetic Classification of Floodplains," Geomorphology 4: 459-486, Nanson, G. C.

and Croke, J. C., 1992.

2.8-13 "Middle and Late Holocene Avulsion History of the River Rhine (Rhine-Meuse Delta, Netherlands)," Geology 22: 711-714, Tornqvist, T., 1994.

2.8-14 "Experimental Study of Avulsion Frequency and Rate of Deposition," Geology 23(4):

356-368, Bryant, M., Falk, P. and C. Paola.

Channel Migration or Diversion 2.8-8

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP I & 2 FukushimaResponse Project Figure 2.8-1 STP Units I & 2 Relative to the Current Colorado River Channel (Dark Blue Line) and Relict Channels of the Colorado River Delta Plain (Red Lines)

ChannelMigration or Diversion 2.8-9

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project EXPLANATION 2

P'.e v- "-w Figure 2.8-2 Quaternary and Tertiary Deposits of the Colorado River, from near Columbus, Texas, to the Gulf of Mexico Channel Migration or Diversion 2.8-10

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STP1 & 2 Fukushima Response Project Source: Modified from Reference 2.8-4 Figure 2.8-3 1838 Map of Texas Showing the Rio Colorado (i.e., the Lower Colorado River) between the Rio-La Vaca and Rio Navidad Rivers to the West and the Rio Brazos to the East ChannelMigration or Diversion 2.8-11

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushiina Response Project 160-Columbus Bend Allo member 3 1401 ------------ Columbus Bend Allo member 2

-"-Eagle Lake Alloformation 120-100-

.0 go-C0 60-40- position of coastal onlap and upstream limits of glacio-eustatic controls 20- on stratigraphic architecture I I I I I I I I I I 250 225 200 175 I50 125 100 75 50 25 Distance Upstream from Modern Shorcline (kin)

Source: Reference 2.8-2. p. 1015 Figure 2.8-4 Longitudinal Profiles for the Lower Colorado River Relative to Mean Sea Level (MSL)

Channel Migration or Diversion 2.8-12

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project 60000 500000 I

Completion of Mansfield Dam i1 J1 2ý tAL111, ilimimiB imiimlllm#alilnmlimiminmmmmmlimimmmmmmm#m I

in 1941

. a..

SEES- Iffin

. h. -.h - ..... a. ....

. -a- '- .. _,_,

._..IIII 0.

-- - . -, S WATER YEAR Note: The flow record includes all historical peaks in the USGS database from water year of 1863 to 2006 Figure 2.8-5 Peak Discharge versus Water Year for the Colorado River at Austin, Texas, (USGS #08158000) before and after the Completion of Mansfield Dam and Lake Travis Channel Migrationor Diversion 2.8-13

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPI & 2 Fukushinta Response Project Source: Figure 2 of Reference 2.8-3 [p. 101]

Figure 2.8-6 Successive Growth Stages of the Modern Delta of the Colorado River, Texas Channel Migration or Diversion 2.8-14

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushinia Response Project

~1 1941 1953 (EST) 7098.22 ACRES 7200 ACRES 6000 5000 1

~1936

/ 4889.88 ACRES

=4000 e j 1933 3470 ACRES 3000 200(

p 1930 1nnr . 1780.46 ACRES 1908 45.8 ACRES n1

, . I - 1

i l I I 1908 1-0 1-5 "0 2 -5 SO 3"5 40 45 5"0 5"5 YEAR S Note: Points marked "P" are postulated for flood years Source: Modified from Figure 4 of Reference 2.8-3 [p. 105]

Figure 2.8-7 Graphic Representation of the Growth of the Colorado River Delta in Acres by Years Channel Migration or Diversion 2.8-15

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushlima Response Project Source: Modified from Fig. 3 of Reference 2.8-3 [p. 103]

Figure 2.8-8 Historical Estuary Occupied by the Colorado River after Abandoning the Caney Creek Area (Solid Line) and the Estuary after Being Filled with Sediments in 1930 (Dashed Line) (Modified from Figure 3 of Reference 2.8-3 [p. 103])

Channel Migration or Diversion 2.8-16

Enclosure NOC-AE-1 3002975 FloodingHazardReevaluation Report STPI & 2 Fukushima Response Project AA' Figure 2.8-9 Plan View of West Branch of the Lower Colorado River, Wild Cow Island, Baxter Island, and McNabb Island Near Matagorda, Texas Channel Migration or Diversion 2.8-17

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP1 & 2 Fukushima Response Project 2.9 Combined Effect Flood Combined effect of different flood causing mechanisms is discussed in Sections 2.1 through 2.8, where applicable. The combined effect flooding criteria for this reevaluation are based on the guidelines presented in ANSI/ANS-2.8-1992 (Reference 2.9-1) and NUREG/CR-7046.

2.9.1 References 2.9-1 "Determining Design Basis Flooding at Power Reactor Sites," ANSI/ANS-2.8-1992, American Nuclear Society, July 1992.

Combined Effects Flood 2.9-1

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushinia Response Project 3 Comparison of Current and Reevaluated Flood Causing Mechanisms Table 3-1 below summarizes the comparison of current and reevaluated maximum flood levels for all possible flood causing mechanisms.

Table 3-1 Current Design Basis and Reevaluation Flood Elevations Difference Reevaluated Bewee FloodMaximum Current maximumBewn Flooding Mechanism Critical Flood Level ft Flood Level ft Reevaluated and Structure Current Flood MSL MSL Levels Local Intense Precipitation Site 32.0 33.0 1.0 ft Flooding in Streams and Site 29.0 26.3 -2.6 ft Rivers 2)

Plant Structures 43.7 36.2 -7.5 ft Upstream Dam Failures ECWIS 39.3 35.8 -3.5 ft Plant Structures 44.5 to 50.8 44.5 to 50.8 0.0 ft MCR Embankment Breach ECWIS 40.8 40.8 0.0 ft Plant Structures 26.74 (3 35.2 (4)

Storm Surge ECWIS 26.74 13) 40.1 (4)

Plant Structures N/A N/A N/A Seiche ECWIS N/A 40.1 N/A Tsunami Site N/A 11.5 N/A Ice Induced Flooding Site N/A N/A N/A Channel Migration or Site N/A N/A N/A Diversion Notes: (1) Reevaluated maximum flood level minus current maximum flood level.

(2) The flood levels are still water levels and do not include wind effects.

(3) Does not include wind-wave effect.

(4) Not compared because the no wind-effect was considered in the UFSAR.

N/A indicates that no impact was identified.

Site includes the plant structures (power block.) and the ECP areas.

Comparisonof Current and Reevaluated Flood Causing Mechanisms 3-1

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 FukushiniaResponse Project 3.1 Local Intense Precipitation The UFSAR prepared for the existing STP I & 2 (Reference 3.1-1) evaluated the effects of local intense precipitation (local PMP) on the safety-related structures.

Probable maximum precipitation for local intense precipitation was taken from HMR 51 (Reference 3.1-1). Point rainfall (10 mi 2 or less) was determined from all-season envelope and was found to be 32.5 in for a 6-hour duration at the STP 1 & 2 site. After losses were estimated and deducted, the total excess rainfall was 31.76 in, reflecting a highly conservative runoff coefficient of 97.7 percent. The rainfall excess was distributed in accordance with the USAGE procedures for determining the standard project flood (Reference 3.1-1).

The two local drainage areas adjacent to the plant structures were considered for the analysis.2 The larger of the two areas lies west and northwest of the plant structures and contains 4.5 mi of land surface. This area drains into the relocated Little Robbins Slough.

A Snyder unit hydrograph was used to develop runoff from the 4.5-mi2 drainage area adjacent to the west side of the plant. Parameters were estimated by analyzing records of rainfall-runoff characteristics of a gauged watershed near the site which has similar hydrometeorological characteristics. The data investigated included seven storms of record which produced approximately one inch of runoff. Each of the events analyzed was reduced to a unit-graph and the seven unit-graphs were averaged. Snyder's parameters were estimated from the average unit-graph, and adjustments were made to these parameters when they were applied to the area under study.

The PMP excess rainfall was applied to the resulting 1-hour unit hydrograph and a peak discharge of 6,400 cfs was calculated. To account for nonlinearity between normal and intense rainfall, the calculated peak was increased by 25 percent, resulting in a peak discharge of 8,000 cfs from the 4.5-mi 2 drainage area and a water level of about 32 ft MSL at the STP 1 &2 site.

The maximum water level was obtained with a HEC-2 hydraulic model that included the relocated Little Robbins Slough and the site.

The other drainage area adjacent to the plant structures lies northeast of the plant and contains about 0.6 mi2 . It drains easterly and southeasterly away from the plant structures through natural streams and into some plant area ditches. The critical point of flow in the discharge of a PMF peak from this area would be at the concrete culvert under the plant access road, just southeast of the ECP, which was designed for a 50-year flood. A local PMP would cause overtopping of the plant access road. Since the elevation of the access road in that area is 30.75 ft MSL for the length of at least 700 linear ft, a water level of 32 ft over a broad-crested weir 700 ft long would pass a discharge of 2,450 cfs assuming a conservatively low value of the weir coefficient C = 2.5. For the 0.6-mi 2 drainage area, this would represent a peak discharge 2 of about 4080 cfs/mi 2, or over two times the peak discharge per mi2 calculated for the 4.5-mi drainage area on the west side of the plant. Therefore, itwas concluded that the water level of 32 ft MSL caused by PMF on Little Robbins Slough would be higher than the PMF on the stream draining the 0.6-mi2 area. However, it was concluded that a PMF on either adjacent drainage area would result in short-term overloading of the plant drainage system but would not enter plant area buildings.

In this reevaluation of local intense precipitation flood hazard the design basis for the local intense precipitation event is the all season one sq. mile or point PMP as obtained from the Local Intense Precipitation 3.1-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPJ & 2 Fukushitna Response Project HRMs 51 and 52. Details of the contributing catchments to the site are provided in Section 2.2.

Peak flood flow discharges were estimated using the U.S. Army Corps of Engineers computer program HEC-HMS and the Natural Resources Conservation Service (NRCS) methodologies.

The hydrologic and hydraulic analyses were performed using the guidelines in NRC NUREG/CR-7046. It was conservatively assumed that all underground storm drains and culverts are clogged and non-functioning during the local PMP storm event. For the purpose of the hydraulic model, the effect of blocked open channels was considered negligible, because the flow conveyed in the channels is small compared to the width of the floodplain during a PMP event. Additionally, the analysis is performed with the conservative assumption that a storm with precipitation depths equivalent to 40% of the PMP occurs prior to the PMP events with 3 to 5 dry days in between the storms as per ANSI /ANS-2.8-1992 (Reference 3.1-2). Water surface elevations were determined using the U.S. Army Corps of Engineers computer program HEC-RAS. Details on the conveyance channels and the vehicular barriers enclosing the power block area are provided in Section 2.2. The maximum water level obtained from HEC-RAS results is approximately 33.0 ft MSL which is greater than the corresponding value of 32.0 ft MSL obtained in the UFSAR for STP 1 & 2. However, the maximum water level from the local PMP event is lower than design basis flood level obtained from the MCR embankment breach analysis.

3.1.1 References 3.1-1. STPEGS Updated Final Safety Analysis Report, Units 1 and 2, Section 2.4, "Hydrologic Engineering," Rev. 15.

3.1-2. "Determining Design Basis Flooding at Power Reactor Sites," ANSI/ANS-2.8-1992, American Nuclear Society, July 1992.

Local Intense Precipitation 3.1-2

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushimna Response Project 3.2 Flooding in Streams and Rivers The PMP values for the PMF analysis in the UFSAR for STP 1 & 2 (Reference 3.1-1) were obtained from HMR-51 and HMR-33, and the spatial distribution was determined based on an extreme historical storm that of June 27 through July 1, 1899 (called the Hearne storm)

(Reference 3.1-1). The temporal distribution of the PMP was determined based on a procedure by USACE (Reference 3.1-1). The PMP was centered at a critical location so that the total rainfall volume is maximized.

The UFSAR for the existing STP 1 & 2 (Reference 3.1-1) evaluated five hydro-meteorologically critical flow scenarios for the Lower Colorado River and selected among these the most critical PMF flow scenario to determine the maximum flood elevation at the STP 1 & 2 site. This study also included a proposed dam at Columbus Bend that was under consideration in the 1960s.

These five PMF flow scenarios are summarized as follows:

The Spillway Design Flood (SDF) for the proposed Columbus Bend Dam, which would result from a PMP storm on the watershed above the dam, was routed to the STP 1 & 2 site. It was assumed that this event would occur in coincidence with the peak of a Standard Project Flood (SPF) from the 755 sq. mi drainage area between the proposed Columbus Bend Reservoir and the STP 1 & 2 site. It was assumed that the peaks of these two floods would be directly additive and that they would occur simultaneously with a base flow of 50,000 cfs. The peak flow at Bay City for this scenario was estimated to be equal to 958,000 cfs (SDF: 648,000 cfs + SPF: 260,000 cfs + base flow: 50,000 cfs) at the STP 1 & 2 site.

" The PMF for the drainage area between Mansfield Dam and Bay City was assumed to occur three days after the occurrence of the SPF over the same area. A base flow of 50,000 cfs was adopted. The peak flow at Bay City for this scenario was estimated to be equal to 913,000 cfs, which includes a base flow of 50,000 cfs.

The SDF outflow hydrograph from Mansfield Dam, which results from the PMF inflow hydrograph into Lake Travis caused by a PMP storm on the watershed above the dam, was routed to the STP 1 & 2 site. This was combined with a SPS occurring over the drainage area between Mansfield Dam and the STP 1 & 2 site, three days after the PMP storm producing the Mansfield Dam SDF. It was also assumed that a base flow of 50,000 cfs occurs simultaneously with the resulting flood. The peak flow for this scenario was estimated to be equal to 698,000 cfs, which includes a base flow of 50,000 cfs.

The PMF for the drainage area between the proposed Columbus Bend Dam and the STP 1 & 2 site was assumed to occur in coincidence with an SPF peak discharge from the proposed dam. It was also assumed that a base flow of 50,000 cfs occurs simultaneously with the resulting flood. The peak PMF for this scenario was estimated to be equal to 894,000 cfs, (PMF: 520,000 cfs + SPF: 324,000 cfs + base flow: 50,000 cfs) at the site.

" A hypothetical PMF for the entire contributing drainage area of the Lower Colorado River basin above the STP 1 & 2 site was assumed, with no credit taken for flood control in the numerous reservoirs upstream from Mansfield Dam, including Lake Travis. The peak PMF for this scenario was estimated to be equal to 1,750,000 cfs.

Flooding in Streams and Rivers 3.2-1

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP I & 2 FukushimnaResponse Project The critical PMF was determined to be the standard design flood (SDF) from proposed Columbus Bend Dam routed to the STP 1 & 2 site and occurring in coincidence with a standard project flood (SPF) on the 755-mi 2 uncontrolled area between Columbus Bend Dam and the STP 1 & 2 site. Addition of a base flow of 50,000 cfs, results in the critical PMF of 958,000 cfs at the STP 1 & 2 site.

The maximum PMF water level at STP 1 & 2 per UFSAR was determined based on the rating curve relationship between Colorado River discharge at Bay City and its stage at the STP Units 1 & 2 site. Therefore conservatively assuming that the PMF discharge of 958,000 cfs occurs at Bay City (instead of the site), the maximum PMF still water elevations was determined to be at elevation of 29 ft MSL (per UFSAR), which is 1 ft above the plant grade elevation of 28 ft MSL.

The flooding resulting from dam failures was found to be more critical than that resulting from the PMF. Therefore, coincident wind-wave activity was considered for flooding resulting from dam failures only.

In this reevaluation of PMF flood hazard, which adopts the PMF analysis conducted for STP 3 &

4 COL Application (Reference 3.2-1) with PMP values obtained from HMRs 51 and 52, the five flood scenarios of possible PMF flows in the Lower Colorado River that were considered for the STP 1 & 2 in the UFSAR were first evaluated for their applicability in determining the maximum flood elevation at the STP site for the present conditions. After careful consideration of the hydro-meteorological setting of the region, it was determined that the five flood scenarios considered for the STP 1 & 2 in the UFSAR cover the permutation of the possible critical flood events that could occur in the region, thus acceptable for the current reevaluation of possible extreme flood conditions.

The first and fourth scenarios considered for STP 1 & 2 in the UFSAR were eliminated because they include the Columbus Bend Dam that was proposed in the 1960s and which met with opposition by different groups at various times. This dam was also referred to later as the Shaw Bend Dam. Plans for the construction of this dam have been abandoned. This was confirmed by conducting an online search, a search of various sources, as well as inquiries to different engineers of the LCRA, none of which revealed any information regarding continuing plans for the construction of the Columbus Bend Dam. The recently published Region "K" Plan for the Lower Colorado Region in the 2007 State Water Plan also states that "Large local opposition to this project was demonstrated at the various Lower Colorado River Water Planning Group (LCRWPG) public meetings and in correspondence during the 2001 LCRWPG plan preparation." The Planning Group's recommendation in the current water plan is to oppose the potential designation of the Shaw Bend site as a potential reservoir site (Reference 3.1-1).

Therefore, it was concluded it is not likely that this dam will be constructed in the future.

The three remaining possible PMF flow scenarios in Lower Colorado River that are analyzed for their effects at STP 1 & 2 are as follows:

(1) The PMF for the drainage area between Mansfield Dam and the Bay City USGS gauging station (3555 sq. mi) combined with an antecedent storm equal to 40% of the PMP occurring over the same drainage area, three days before the PMF. This combined flow is added to the flow release from Mansfield Dam and to the base flow at Bay City to determine the peak PMF flow for this scenario.

Floodingin Streams and Rivers 3.2-2

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project (2) The PMF inflow hydrograph to Mansfield Dam, which results from a PMP storm on the watershed upstream of the dam (from Lake O.H. Ivie to Mansfield Dam), routed through Lake Travis and combined with the flood hydrograph from a sequential storm equal to 40% of the PMP occurring over the drainage area between Mansfield Dam and Bay City (3555 sq. mi), three days after the PMP storm upstream of Mansfield Dam. This combined flow is added to the base flow at Bay City to determine the peak PMF flow for this scenario.

(3) The PMF for the Lower Colorado River basin area between Lake O.H. Ivie and Bay City (18,197 sq. mi) combined with the flood hydrograph from an antecedent storm equal to the SPS over the same area, occurring three days before the PMF. This combined flow is added to the base flow at Bay City to determine the peak PMF flow for this scenario.

Conservatively, this scenario does not account for the storage effect of Lake Travis at Mansfield Dam nor any other dam in the Lower Colorado River basin.

From the three scenarios considered above, Scenario 1 produces the highest peak PMF at Bay City. The highest flood peak at Bay City is caused by the PMP for the drainage area between Mansfield Dam and the Bay City combined with an antecedent storm equal to 40% of the PMP occurring over the same drainage area, the flow release of 90,000 cfs from Mansfield Dam, and the base flow of 5200 cfs. Therefore, the peak flow of 1,397,432 cfs for Scenario 1 is used as the most critical PMF scenario to evaluate potential flooding at the STP 1 & 2 site in this reevaluation.

The maximum still water surface elevation at the STP 1 & 2 site for the peak PMF discharge of 1,397,432 cfs was estimated using the United States Army Corps of Engineer's HEC-RAS hydraulic model (Reference 3.1-1). The HEC-RAS model for the STP 1 &2 site was developed on the basis of topographic data and hydraulic characteristics such as Manning's roughness coefficients.

The maximum PMF still water surface elevation at the STP 1 & 2 site assuming a normal depth condition at the downstream boundary was estimated to be equal to 26.1 ft NAVD88 (26.3 ft NGVD29 or MSL) in this reevaluation, which is lower than the plant grade elevation of 28 ft MSL.

According to the FSAR COLA STP 3 &4 (Reference 3.2-1), the flooding resulting from dam failures was found to be more critical than that resulting from the PMF, same as the finding from the UFSAR for STP 1 &2. Therefore, coincident wind-wave activity was considered for flooding resulting from dam failures only.

To summarize, the maximum PMF still water level of 26.3 ft NGVD29 (or MSL) at the STP 1 & 2 site from this reevaluation is lower than the corresponding value of 29 ft MSL given in UFSAR for STP 1 & 2.

3.2.1 References 3.2-1 South Texas Project Units 3 & 4 Combined Licensing Application (COLA), Final Safety Analysis Report (FSAR), Rev. 7, Nuclear Innovation North America LLC, February 1,2012.

Floodingin Streams and Rivers 3.2-3

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushitna Response Project 3.3 Dam Breaches and Failures This section is divided into two parts. The first part deals with upstream dam failures and the second part deals with the MCR embankment breach.

3.3.1 Dam Failures The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated flooding which could result from postulated failures of dams upstream on the Colorado River. The major dams considered were two existing dams (Mansfield and Buchanan) and one proposed dam at the time (Columbus Bend). Two dam failure sequences and one cascade dam failure of all upstream dams were analyzed. Details of the three cases are given below.

(1) Buchanan Dam was assumed to fail with the reservoir level at El. 1020.5 ft MSL and the resultant discharge forms the inflow hydrograph to Lake Travis with that reservoir assumed at an initial elevation corresponding to the SPF. The discharge hydrograph of the Buchanan failure was initially reduced by the storage volume in Lake Travis between the SPF elevation and top of dam El. 750.1 ft MSL and the remainder of the Buchanan failure hydrograph was assumed to be the inflow hydrograph to Lake Travis during the instantaneous failure of Mansfield Dam. A peak discharge of about 9 million cfs was conservatively estimated following the failure of Mansfield Dam. The volume of the dam-break hydrograph was about 4.5 million acre-feet which is in excess of the combined storage of Buchanan Dam, Mansfield Dam, and three smaller dams between them. As an initial condition, the river was assumed to be at the stage corresponding to a flow equal to the SPF for the uncontrolled area above the Bay City gauge. Based on the initial conditions assumed, the discharge resulting from the failure of Mansfield Dam was routed downstream to the Bay City USGS gauge using the NWS Dam-Break Program.

Final plant water surface elevations were obtained from the stage-discharge relationship curve. The peak discharge at Bay City was found to be about 1.9 million cfs. This would result in a maximum water surface elevation of about 32 ft MSL at the site. The peak discharge at Bay City was found to occur 65 hours7.523148e-4 days 0.0181 hours 1.074735e-4 weeks 2.47325e-5 months after the failure of Mansfield Dam and therefore it concluded that there is sufficient time to implement emergency procedures at the STP 1 & 2 site.

(2) The proposed Columbus Bend Dam was assumed to fail in series with the above mentioned Mansfield Dam failure. The initial level of Columbus Bend Reservoir was assumed to correspond to that for the SPF in the Colorado River; Columbus Dam was assumed to fail when flood waters from Mansfield Dam reached the top of Columbus Dam. This analysis indicates a peak discharge of about 1.8 million cfs would result at Bay City 67 hours7.75463e-4 days 0.0186 hours 1.107804e-4 weeks 2.54935e-5 months after the failure of Mansfield Dam. This would correspond to a static water elevation of 31.7 ft MSL at the STP 1 &2 site. Thus, the postulated Mansfield Dam failure prior to construction of Columbus Bend Dam was the critical failure sequence.

(3) All dams above Columbus Bend were conservatively assumed to fail and release their top-of-dam contents into the river. The failure hydrographs were assumed to arrive as Mansfield Dam with such timing as to maximize the breach outflow hydrograph from Mansfield. The antecedent flow between Mansfield and Bay City was conservatively taken to be the 958,000 cfs derived from the Columbus Bend SDF and the SPF on the area between Columbus Bend and the STP 1 & 2 site. Although this implies that Columbus Bend would be in place, no credit was taken for any possible attenuation it Dam Breaches and Failures 3.3-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPI & 2 Fahkushima Response Project might produce. A 600-foot-wide by 200-foot-high section of Mansfield Dam was assumed to fail when overtopped by five feet of water, producing an instantaneous peak outflow of 5.5 million cfs. Attenuation over the estimated 238 flow miles between Mansfield and Bay City reduced the peak flow to about 3.4 million cfs resulting in a still water level at the STP site of 34.1 ft MSL.

From the three scenarios considered, the third scenario results in the maximum still water level (34.1 ft MSL). With the consideration of 2-year winds, the maximum water level including wind-wave runup was estimated to be 43.7 ft MSL at all STP 1 & 2 structures except the ECWIS. For the ECWIS, the maximum water level was estimated to be 38.9 ft MSL. However, the maximum water level for ECWIS, including wind setup and runup was estimated to be 39.3 ft MSL for the first dam failure scenario in which Buchannan and Mansfield Dams are assumed to fail.

In this reevaluation of PMF hazard, which adopts the PMF analysis conducted for STP 3 &4 COL Application (Reference 3.2-1), two dam failure scenarios were considered as detailed below.

(1) Simultaneous failure of all upstream dams induced by a seismic event. The failure is to occur coincidentally with a 2-year design wind event and a 500-year flood or a one-half PMF.

(2) Domino-type failure of upstream dams with the same coincidental wind and flood events as in first scenario. It is postulated that the upstream-most dam(s) would fail first, thereby releasing a dam break flood wave (or waves) that propagates downstream and triggers the failure of the downstream dams one after another in a cascading manner. It is assumed that the 56 dams on the Colorado River and its tributaries upstream of Buchanan Dam (with top-of-dam capacity over 5000 AF) would fail in such a manner that their flood flow, expressed in terms of their respective top-of-dam storage volumes, would arrive at Lake Buchanan at approximately the same time, triggering the failure of Buchanan Dam. The dam break flood flow from Buchanan Dam would then propagate downstream to Lake Travis, overtopping Mansfield Dam and causing it to fail. The dam break flood from Mansfield Dam then propagates downstream to the STP 1 &2 site.

Of these two scenarios, the second scenario would generate the most critical flood level at STP 1 & 2 because of the deliberate alignment of the travel and arrival of the dam breach flood volumes and flood peaks from the major upstream dams. Therefore, only the flood risk resulting from the second scenario was evaluated further as detailed below.

In the conceptual dam break flood model, the 56 dams upstream of Buchanan Dam would fail in a domino manner, with their combined top-of-dam storage capacity, totaling 6.87 MAF, arriving at Buchanan Dam at approximately the same time. As the flood level at Buchanan Dam rises to about 3 ft over the dam crest elevation of 1025.35 ft MSL, the dam would fail, thereby releasing the flood storage of Buchanan Dam plus the combined flood volumes from the 56 upstream dams. In accordance with the combined events requirements stipulated in the American National Standard ANSI/ANS-2.8 (Reference 3.1-2), the evaluation of potential flood risks as a result of non-hydrologic dam break failures should also consider a coincidental event equal to a 500-year flood or one-half PMF, whichever is less. In this analysis, a constant flood flow of 500,000 cfs, slightly higher than the peak Standard Project Flood (SPF) inflow at Buchanan Dam and the 500-year flood peak inflow at Mansfield Dam, was conservatively used to represent the coincidental flow. The 500,000 cfs coincidental flow was applied to the entire Dam Breaches and Failures 3.3-2

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushimna Response Project model reach from Buchanan Dam to the downstream boundary at 4600 ft (0.9 river miles) upstream of the Gulf Intracoastal Waterway. The flood wave from the breaching of Buchanan Dam would propagate down to the 266.4-ft high Mansfield Dam, with a crest elevation at 754.1 ft MSL and a top-of-dam storage capacity of 3.30 MAF. Mansfield Dam was postulated to fail when it was overtopped by 3 ft at El. 757.1 ft MSL. The three dams located between Buchanan and Mansfield Dams: Roy Inks, Alvin Wirtz, and Max Starcke Dams, have a combined storage of about 298,300 AF. These dams were not assumed to fail in the dam break model because their combined total storage amounts to only about 9% of the total dam break flood volume at Mansfield. The SPF flood hydrographs from 19 tributaries between Buchanan and Mansfield Dams were included as tributary inflows to this reach. The tributary inflows together with the dam break flood wave from Mansfield Dam were then routed to the STP 1 & 2 site in the HEC-RAS model, which predicted maximum still water level of 28.6 ft MSL with flood wave travel time of 65 hours7.523148e-4 days 0.0181 hours 1.074735e-4 weeks 2.47325e-5 months after the failure of the Mansfield Dam.

Wind setup and wind-wave runup estimated for STP 3 & 4 are not applicable to STP 1 & 2 because of difference in plant grade. For STP 3 & 4, the power block area is filled up to 34 ft MSL whereas STP 1 & 2 structures are built on grade with nominal elevation of 28 ft MSL.

Consequently, additional evaluation was made to estimate the wind setup and wind-wave runup on STP 1 & 2 based on the still water level obtained for STP 3 & 4. Accordingly, the maximum water levels at the plant structures and ECWIS are approximately 36.2 ft MSL and 35.8 ft MSL, respectively. The corresponding maximum water levels from the UFSAR for STP 1 & 2 are approximately 43.7 ft MSL and 39.3 ft MSL, respectively.

3.3.2 MCR Embankment Breach The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated flooding which could result from postulated failure of MCR embankment.

Even though the postulated instantaneous failure of the MCR embankment is not a credible event and postulation of such an event reflects a conservative consideration, such a failure was analyzed. The conservatisms included in the analysis and results are presented below.

(1) The engineering design of the MCR embankment for protection against seepage and erosion incorporates conservative measures which result in the conclusion that it is extremely unlikely that a failure of the embankment might occur as a result of seepage or erosion.

(2) Historical behavior of rolled-earthfill embankment in earthquakes which are from two to four times the magnitude of the STP 1& 2 Safe Shutdown Earthquake (SSE), as well as results from site soil investigations, indicates that it is extremely unlikely that the MCR will fail due to a seismic event the equivalent of the STP 1 & 2 SSE.

(3) The most conservative approximation of a postulated failure suggested by noted soil mechanics authorities would be the translation of a 600-ft length of embankment approximately 200 ft downstream. Further conservative approximations equate the resulting opening to a 400-foot-wide rectangular opening, whereas the STP 1 & 2 analysis considers an opening of approximately 2,000 ft.

(4) Although historical failures involve a time-related rate of opening, the conservatism of an instantaneous removal is adopted to determine flood levels.

Dain Breaches and Failures 3.3-3

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushima Response Project (5) The embankment breach analysis for STP 1 & 2 takes no credit for flow retardation and dispersion that, in reality, would be provided by the circulating water intake pipes, Circulating Water intake Structure (CWlS), Circulating Water Discharge Structure (CWDS), and various other obstructions between the embankment and plant structures.

The analysis of the reservoir embankment failure was conducted by using a two dimensional computer program entitled "System 21" developed by Danish Hydraulic Institute (DHI), and the one-dimensional NWS Dam-Break Program. Three different models were used in this study as discussed below.

(1) The Danish Hydraulic Institute's "System 21" model was used to simulate flood-wave impacts on the south side of the power block and the ECW Intake Structure. It fell slightly short of producing the highest water levels, but it did demonstrate the relationship between breach length and water level as well as the fact that quasi-steady-state computations would be appropriate for detailed evaluation of water levels around the site.

(2) The "System 21"model was used in a steady-state mode to assess the water levels which would be expected several minutes after the postulated breach. The results from this model produced the highest water levels, which were used for design purposes.

(3) The NWS Dam-Break Program was used to assess the flood-wave impact on the southern ECP embankment.

The results of the analysis indicate that the maximum water levels at the plant structures (power block) to vary between 44.5 and 50.8 ft MSL and at the ECWIS to be 40.8 ft MSL.

This reevaluation of the MCR embankment breach flooding hazard adopts the approach and methodology for estimating the flood wave generated by an embankment breach in the STP 3 &

4 COLA analyses (Reference 3.3-1). In addition, an independent MCR breach analysis conducted for STP 3 & 4 and also a 2012 MCR embankment breach analysis made for STP 1 &

2 to determine the impact of debris from the embankment breach (Reference 3.3-2) were used.

Review of all the analyses indicates that the analysis presented in UFSAR for STP 1 & 2 (Reference 3.1-1) still provides a bounding analysis and meets the objectives of the flooding reevaluation of using present-day methodology and data. This includes flooding water level and associated forces on structures, effects of scouring, erosion, sedimentation, waterborne missiles and debris. In conclusion, the assessment shows that the current design basis, which is based on the MCR breach flooding hazard, will not be exceeded and that there is a margin available than what is presented in the UFSAR for STP 1 & 2.

3.3.3 References 3.3-1 South Texas Project Units 3 &4 Combined License Application (COLA), Final Safety Analysis Report (FSAR), Rev. 7, Subsection 2.4S.4 (Potential Dam Failures),

Nuclear Innovation North America LLC (NINA), February 1, 2012.

3.3-2 South Texas Project Units 1 &2 Flood Analysis, prepared by Atkins, Austin, Texas for STPEGS, Document No. 120021, March 2012.

Dain Breaches and Failures 3.3-4

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 3.4 Storm Surge The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated potential flooding resulting from probable maximum surge due to probable maximum hurricane (PMH) winds. The method and numerical values for PMH parameters (central pressure index, the radius of maximum wind, the hurricane translation speed, and the peripheral pressure) were based on NOAA's Technical Report NWS23 (Reference 3.4-1). Because both the radius of maximum wind and the forward speed of storm center are variable, the maximum surge elevation caused by the PMH four combinations of radius of maximum wind and forward speed were considered in the analysis. These were obtained by combining the extreme values of the radius of maximum wind and the forward speed as indicated in Table 3-2.

In calculating the hurricane surge in the UFSAR (Reference 3.1-1), a computer program entitled, "Storm Surge on The Open Coast - Fundamentals and Simplified Prediction",

developed by the USACE Coastal Engineering Research Center, was utilized. In order to support the results obtained from the analysis, the bottom friction value was calibrated and verified. This was accomplished by applying the above-referenced program along the traverses of historical hurricanes in the vicinity, and using the data and physical characteristics of these hurricanes to reproduce the historical surge peak as nearly as possible. The selected hurricanes are Hurricane Carla (1961) and Hurricane Celia (1970). Pertinent data for these two hurricanes were obtained from NOAA's "The Monthly Weather Review".

To estimate the maximum surge at the coast, the four combinations of the radius of maximum wind and translation speed as shown in Table 3-1 were analyzed. The initial rise of water surface 2.5 ft was specified based and the astronomical high tide of 2 ft above mean low water (MLW) was assumed to occur in coincidence with the hurricane surge. A 10-percent increase in Van Dorn's wind stress was used to account for the additional stress caused by energy imparted to the sea due to precipitation. A computational traverse was drawn perpendicular to the bottom contours on a line extending into the Gulf from a point near the mouth of the Colorado River. The PMH was assumed to move from the Continental Shelf, along the west side of the traverse, toward the coast. The center of the hurricane was placed at a distance equal to the radius of maximum wind from the traverse line.

The maximum surge elevation at the mouth of the Colorado River was calculated to be El. 25.08 ft MSL. It was emphasized that RG 1.59 gives an estimated probable maximum surge El. of 23.5 ft MLW (22.07 ft MSL) at Freeport on the Gulf Coast near the STP 1 & 2 site. Therefore, it was concluded that the computed maximum surge elevation of 25.08 ft MSL is conservative.

To determine water surface elevation at the STP 1 & 2 site resulting from the maximum PMH surge concurrent with a 100-year flood in the Colorado River, a backwater profile was run using the "HEC-2 Water Surface Profile" program. The water surface elevation at the site was determined to be 26.74 ft MSL, which is below the plant grade elevation of 28.0 ft MSL.

Therefore, the PMH was not considered to be a design basis event for maximum water surface elevation and hydraulic forces at the plant structures. In addition, wave action was not considered because the water surface elevation resulting from the PMH was below the plant grade and was concluded that would be no resonance effect at the STP 1 & 2 site area.

In this reevaluation of probable maximum storm surge (PMSS) flooding hazard, similar to that in the UFSAR, the PMH would produce the probable maximum meteorological wind (PMMW) at Storm Surge 3.4-1

Enclosure NOC-AE-13002975 Flooding Hazard Reevaluation Report STPI & 2 Fuhkushima Response Project the STP 1 & 2 site. The PMH parameters were obtained from the NOAA Technical Report NWS23 (Reference 3.4-1) based on the milepost location along the Gulf Coast.

The coupled ADCIRC two-dimensional depth-integrated model (ADCIRC-2DDI, version 49) and the SWAN wave model were developed for the STP 3 & 4 COLA (STP 3 & 4 COLA ADCIRC model). The PMH parameters (Ap, radius of maximum wind, forward speed, track direction),

were used to define the physical attributes of the PMH in the model. Model simulations are performed with numerous combinations of input PMH parameters to obtain the PMSS elevation.

The effect of wind-wave run-up was superimposed on the PMSS elevation to obtain the maximum water level at the STP facilities. As defined in Regulatory Guide 1.59, the antecedent water level includes the 10 percent exceedance high spring tide, initial rise, and the long-term rise. Combining the individual contributions, an antecedent water level of 5.1 feet MSL (4.9 ft NAVD 88) was used in the STP 3 &4 COLA ADCIRC model.

A total of 88 model sensitivity runs were performed for hurricane scenarios combining various PMH parameters. A sensitivity analysis was also performed for the choice of the wind model to be used in STP 3 & 4 COLA ADCIRC. Based on the results from the sensitivity cases, the PMSS elevation at the STP site was obtained as 29.3 ft MSL. This PMSS will occur as the result of a hurricane traveling towards northwest direction (i.e., an approach direction of 1350 clockwise from the north) passing within 24 miles of the STP site. During its life up to the point of landfall, the storm will have a constant forward speed of 23 miles per hour, a central barometric pressure of 887 millibar, and a maximum sustained wind speed (1-min average) of 160 knots (184 miles per hour).

Wind wave rununp on STP 1 & 2 safety-related structures were evaluated separately. As the area south and southeast of the ECP would be inundated during the PMSS event, the waves affecting the ECP embankment would be limited by the PMH wind duration. With the PMH forward speed after landfall at about 20 mph and the site located approximately 16 miles from the Gulf of Mexico shoreline, the maximum wind would be sustained for less than one hour.

Considering a duration limited wave for a maximum period of one hour, the deep water significant wave height and peak period were estimated at 13.9 ft and 4.9 seconds, respectively.

The wave height would reduce due to wave breaking before reaching the ECP embankment because of the limited water depth over the grade elevation of about 27 ft MSL at the foot of the ECP embankment. The water depth of 2.3 ft at the exterior toe of the ECP embankment would sustain a significant and maximum breaking wave height of approximately 1.4 and 1.9 ft, respectively. The significant, 2 percent, and maximum wave runup on the ECP embankment, which has a crest elevation of 34 ft MSL and 1V:3H side slope, were estimated at 3.2 ft, 4.4 ft and 5.9 ft, respectively. The storm surge level including the significant, 2 percent and maximum wave runup on the ECP embankment would be 32.5 ft MSL, 33.7 ft MSL and 35.2 ft MSL, respectively. While the water level associated with the significant and 2 percent wave runup would remain below the ECP embankment crest, the water level associated with the maximum wave runup would exceed the ECP embankment crest elevation. However, because of the 4.7 ft freeboard, defined as the height of the ECP embankment crest above the PMSS still water level, the overtopping rate on the ECP embankment for the duration of the PMSS would be very small and would have negligible effect on the water level inside the ECP. The maximum runup level at the ECWIS as a result of PMH wind action, in combination with a 100-year 4-day precipitation, is predicted to be 40.1 ft MSL as part of seiche flooding evaluation described in Section 2.5.

Storm Surge 3.4-2

Enclosure NOC-AE-13002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project The cooling water discharge structure is located on the west ECP embankment as the ECWIS but on the north side of the dividing dike. It has a lower profile than the ECWIS, with the top elevation at 36 ft MSL. The 40.1 ft MSL maximum wave runup level at the ECWIS will provide a bounding flood level for the discharge structure as well during a PMH event. The maximum reevaluated flood elevation including wave runup at the ECP embankment exterior as 35.2 ft MSL could also conservatively be adopted for the power block safety-related SSCs.

3.4.1 References 3.4-1 Schwerdt, R. W., Meteorological Criteria for Standard Project Hurricane and Probable Maximum Hurricane Windfields, Gulf and East Coast of the United States, NOAA Technical Report NWS23, National Weather Service, Sept. 1979.

Storm Surge 3.4-3

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP1 & 2 Fukushima Response Project Table 3-2 Probable Maximum Hurricanes Considered in the UFSAR for STP 1 & 2 (taken from Reference 1.1-3)

Radius of Central Peripheral Maximum Maxiumn Maxtimum Pressure Pressure Wind (R) Translation Gradient Wind Sustained Case Index (CPI) (Pn) (nautical Speed (Vgx) Speed Wind Speed No. (inches) (inches) miles) (knots) (Vx) (knots)

(knots) 1 26.16 30.12 136.9 134.7 2 26.16 30.12 5 20 136.9 140.0 3 26.16 30.12 21 134.9 132.8 4 26.16 3012 21 20 134.9 138.1 Storm Surge 3.4-4

Enclosure NOC-AE-1 3002975 Flooding HazardReevaluation Report STP I & 2 Fukushina Response Project 3.5 Seiche The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated potential flooding from seiches and concluded that the flooding at the site due to seiche effect is considered insignificant because seiche has not been considered as the controlling influence for Gulf of Mexico and Matagorda Bay, the large water bodies in the immediate vicinity of the site.

In this reevaluation of seiche flooding hazard, the impact of seiche initiated by tsunami, seismic, atmospheric and PMH conditions are considered. The reevaluation considers the water bodies at the site, the ECP and MCR, and also the large water bodies at the immediate vicinity of the site, Gulf of Mexico and Matagorda Bay.

For the ECP, the PMH wind condition is found to be critical in producing the maximum water level at the ECW intake structure. The PMH wind condition was combined with 100-year, 4-day rainfall event (1.3 ft) per the guideline provided in the ANSI/ANS-2.8-1992 for considering combined events (Reference 3.1-2). Therefore, for the combined PMH and the ECP 100-year water level event, the maximum wind setup of approximately 0.6 ft, resulting in a maximum still water elevation of about 27.9 ft MSL is estimated. For this event, the significant and maximum wave heights and peak wave period, at the toe of the ECP embankment, are determined to be 1.3 ft, 1.7 ft and 2.3 sec, respectively. Such wave generates a maximum wave runup of about 12.2 ft, resulting in a maximum water elevation of about 40.1 ft MSL at the ECWIS based on the conservative assumption of direct wave attack angle.

The natural period of the ECP was estimated to range between 3.7 and 4.7 min. with a depth ranging between 8.6 ft and 14 ft, respectively, corresponding to the minimum operating water level and the combined maximum operating water level and the PMP. Consequently, because of the excitation periods of potential tsunamis, seismic, atmospheric and PMH conditions are significantly different than the natural period of the ECP, flooding of the STP 1 & 2 site due seiche is precluded. In addition, there is no seiche flooding potential from MCR as described in Reference 3.2-1 (STP 3 & 4 FSAR COLA, Subsection 2.4S.8).

At the Gulf of Mexico and Matagorda Bay, seiche is not considered to be a controlling influence to pose a flooding hazard on the STP 1 & 2, because of the considerable distance between the water bodies and STP 1 & 2 site. Other than for floods on the Colorado River, the hurricane storm surge is the dominant factor responsible for coastal area flooding. Therefore, the flooding at the site due to seiche effects is considered insignificant.

Seiche 3.5-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP I & 2 Fukushima Response Project 3.6 Tsunami The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated historical and potential flooding which could result from tsunami and concluded that no consideration of the effects of tsunami on safety-related facilities is warranted.

Historically reported earthquakes, which could reasonably be expected to have affected the STP 1 & 2 site area were reviewed, however, no record was found of any tsunami occurrences in the vicinity of STP 1 & 2 site that could have resulted from the earthquakes.

The geoseismic potential of the coastal area in the vicinity of STP 1 & 2 was investigated as a possible generating source of a tsunami. Evaluation of potential tectonic faults in the coastal region near the site indicated that there is no known source that might generate such a fault.

The coastal plain consists of flat, gently sloping coastal terraces, approximately 50 to 75 miles wide. Therefore, there is no dominant physical relief to precipitate potential landslides of any significant magnitude. In addition, there is also no active volcanism in the coastal plains area that could generate potential tsunami. In view of the foregoing, no further consideration of the probable maximum tsunami water level was found to be warranted.

This reevaluation of the tsunami flooding for STP 1 & 2 adopts the same approach and methodology as the STP 3 & 4 COLA FSAR evaluation (Reference 3.2-1) which was prepared in 2008, and supplements with recent data from literature and databases. The COLA evaluation followed the hierarchical approach described in NUREG/CR 6966. In addition, the COLA evaluation is consistent with the requirement of NUREG/CR-7046 which specifies that the antecedent water level of 10% exceedance high tide and the effect of sea level rise should be considered for tsunami flooding hazard assessment for coastal sites. New information regarding the historical tsunamis, source mechanisms, and propagation of tsunamis in the Atlantic Ocean along the U.S. coast was examined. The result of the examination suggests that the tsunami sources and amplitudes used as a basis for the modeling effort in COLA remain valid for the flooding reevaluation of STP 1 & 2 safety-related facilities. It was postulated that the probable maximum tsunami (PMT) source would be submarine mass failure of the East Breaks slump.

The study presented in the COLA predicts a maximum tsunami runup level of 11.5 ft MSL (including 10% high tide and 100-year sea level rise) at the coastal area near the site for the PMT, while a study by USGS (Reference 3.6-1) predicts a corresponding maximum water level of 16.5 ft MSL under more conservative physical assumptions.

3.6.1 References 3.6-1 "Regional Assessment of Tsunami Potential in the Gulf of Mexico: U.S. Geological Survey Administrative Report. September 2, 2009.

Tsunani 3.6-1

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP I & 2 Fukushima Response Project 3.7 Ice Induced Flooding The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated ice effects based on a review of historical river water temperature data for the Colorado River which indicated that there is a general trend of an increase in water temperature as the river flows south. A research of literature revealed that no record of historical ice flooding in the area of the plant site.

Sufficient information for a complete analysis of river temperatures near STP 1 & 2 site could not be obtained as only random observations were available for irregular intervals.

Consequently, the river water temperature data at the USGS gauge at Wharton, Texas, was used. Because the USGS gauge at Wharton is upstream of the USGS gauge at Bay City, the values used reflected lower temperatures and a more conservative estimate relating to the probability of ice produced flooding at the STP 1 & 2 site. The data covered October 1944 through September 1975 with some gaps and the lowest river water temperature recorded was 35°F, which occurred on December 23, 1963, and on January 14, 1964.

A probability analysis of the annual minimum water temperature showed that the minimum river water temperature of 32°F has a probability of occurrence of one in every 10,000 years.

Because of the low probability of a 32°F water temperature, and because the plant site is adjacent to the reach of the Colorado River which is subject to tidal effects, it was concluded that ice flooding is not a potential hazard at the STP 1 & 2 site.

In this reevaluation of ice induced flood hazard, which adopts the analysis conducted for STP 3

& 4 COL Application (Reference 3.2-1), the effects of low temperature and ice formation on the flooding hazard of the STP 3 & 4 safety-related facilities were evaluated using historical data up to 2006, which showed that low air and water temperature or ice jam events in the Lower Colorado River would not cause any risk of flooding that would challenge the safety functions of the Units 3 & 4 plant. The ice induced flooding reevaluation for Units 1 & 2 adopts the Units 3 &

4 analysis, augmented by an examination of the most recent data from 2006 to 2012.

Specifically, the recent data (including the daily extreme air and water temperature, and ice jam data) from 2006 to 2012 were compared with those from the STP 3 & 4 analysis (Reference 3.2-

1) to confirm that the conclusions from the STP 3 & 4 analysis are still bounding and valid for STP 1 & 2. Examination of the most recent air and water temperature data and a search of the ice jam database support the conclusions of STP 3 & 4 COL Application that significant ice formation in the area of the STP site and in the Colorado River upstream of the site that may affect flooding of STP 1 & 2 safety related facilities and functions is highly unlikely.

Ice Induced Flooding 3.7-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STP1 & 2 Fukushima Response Project 3.8 Channel Migration or Diversion The UFSAR prepared for the existing STP 1 & 2 (Reference 3.1-1) evaluated historical and potential channel migration or diversion and concluded that because of flood regulation by upstream reservoirs and the responsibility for channel stabilization and improvement delegated by Congress to USACE, channel diversion is not considered to be a significant factor to the safety of the STP 1 & 2 site.

In this reevaluation of channel migration or diversion, which adopts the analysis conducted for STP 3 & 4 COL Application (Reference 3.2-1), review of the evidences and potential causes of channel diversions in the Colorado River were made which indicates that there is little likelihood that major channel diversions impacting the safety facilities and function of STP Units 1 & 2 would occur. Specifically, flooding events in the order of a PMF at the STP 1 & 2 site as a result of channel diversions is considered improbable. Similarly, interruption of the non-safety water supply to the STP Reservoir Makeup Pumping Facility located on the west bank of the Colorado River as a result of channel migration is considered unlikely.

Channel Migration or Diversion 3.8-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPI & 2 Fukushinia Response Project 3.9 Combined Effect Flood Combined effect of different flood causing mechanisms is discussed in Sections 3.1 through 3.8, where applicable.

CombinedEffect Flood 3.9-1

Enclosure NOC-AE-1 3002975 Flooding Hazard Reevaluation Report STP I & 2 Fukushima Response Project 4 Interim Evaluation and Actions Taken or Planned According to the 10 CFR 50.54(f) Request For Information dated March 12, 2012, licensees whose current design basis does not bound the reevaluated flooding hazard are also requested, within 2 years of submitting the Flooding Hazard Reevaluation Report, to provide an Integrated Assessment Report that evaluates the total plant response to the reevaluated hazard. In addition, the licensees performing an integrated assessment are also requested to provide a list of any interim actions, taken or planned, to address the reevaluated hazard while the longer-term analysis takes place. For STP Units 1 & 2, no interim actions or integrated assessment are necessary as explained below.

For STP Units 1 & 2 plant, it is determined that, based on the results of the reevaluation of each of the applicable flooding mechanisms for the site, the current design basis flood protection measures implemented at the site will provide adequate protection against the reevaluated flood hazards. Specifically, as described in Section 3, the flooding reevaluation confirms that the controlling flooding mechanism in the current design basis for the plant structures and the ECWIS remains to be the postulated instantaneous MCR embankment breach. Further, the results of the flooding reevaluation provided in Section 2, demonstrate that the current design basis flood elevations which vary between 44.5 and 50.8 ft MSL for the plant structures (power block) and 40.8 ft MSL for the ECWIS remain to be the bounding water levels from all external event flood sources. In addition to the maximum flood elevations, the reevaluation also includes, as appropriate, assessments of hazards associated with flooding such as duration of inundation, impact forces on structures, effects of scouring, erosion, sedimentation, waterborne missiles and debris.

No adverse impacts have been identified as described in Section 2, even though the flood levels provided in the STP Units 1 & 2 UFSAR for four (4) of the non-controlling flooding mechanisms, local intense precipitation, storm surge, seiche and tsunami, are not bounded by the corresponding reevaluated flood levels.

During the performance of flooding walkdowns and compilation of the flooding walkdown report, an initial identification of available physical margin (APM) was made. An item which would cause adverse impact to safety related SSCs if exposed to external flood waters (such as an unsealed exterior penetration or non-watertight door) was considered to have small APM when its minimum opening level or flood protection level is less than one foot above the licensing basis flood height at that location. The flooding walkdown report submitted in response to NTTF Recommendation 2.3 committed to address the related effects of small APM during the NTTF Recommendation 2.1 activities. As indicated in Table 3-1, the re-evaluated flood level for each of the flooding events (including combination events) is less than the current governing design basis flood level for every critical structure. Thus, there is no risk significance associated with the items identified with small APM, as the items remain above the reevaluated flood levels.

Interim Evaluations and Actions Taken or Planned 4-1

Enclosure NOC-AE-1 3002975 FloodingHazard Reevaluation Report STPI & 2 FukushiniaResponse Project 5 Additional Actions No additional action required.

Additional Actions 5-1

ML13079A808

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