Text

PMF Near Midland Site, Submitted as Part of Amend 31 to PSAR
	ML19329F064
Person / Time
Site:	Midland
Issue date:	12/31/1975
From:	; BECHTEL GROUP, INC.
To:	;
References
	NUDOCS 8006190861
	Download: ML19329F064 (41)
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, Consumers Power Company Midland Units 1 & 2 Job 7220 i

PROBABLE MAXIMUM FLOODING NEAR THE MIDLAND SITE i

7 THis DOCUMENT CONTAINS P00R QUAUTY PAGES 3 l

1 BECHTEL December 1975 v (Revision 3)

,y- . -- e v --. ., , - -- ,. , n , ,_m,. -. -,-*,m.--------.-em-, -- --

t PROBABLE MAXIMUM FLOODING NEAR THE MIDLAND SITE

1. INTRODUCTION The main hydrologic concern regarding the safety of vital installations is the probable maximum water level at the 1 plant site. Five parts are involved in this problem: 3

1. What would be the natural probable maximum flood (PMF) discharge in the Tittabawassee River?

2. What would be the effect of upstream dam fail-ures at the time of the PMF?

3. What would be the maximum water level resulting from the PMF (or PMF plus dam failure) ?

4. What design features are necessary to protect the plant against concurrent wind activity? 1 I

S. Could a Bullock Creek PMF with a 100-year flood on the River give a higher water level?

2. CONCLUSION The Midland Plant's safety-related facilities will be ,1 immune from hazards associated with probable maximum flooding. This immunity derives from:

. setting the plant 3 feet above the probable maximum flood level.

. providing for sand-bag protection of vital entrances against possible wave run-up This report presents tlie details of the study upon which this conclusion is based.

3. PREVIOUS STUDIEQ Brater and Wisler (1952) have determined a probable maximum flood for the Tittabawassee River just above 3 the Midland power plant site.

f Their calculations at that time did not provide a rigorious analysis of the effects of failure of the upstrea% dams due to overtopping. They estimated this effect by applying the dams' contents as equi-valent rainfall to the basin unit hydrograph. This was admittedly a rough approximation, but the contents are relatively small, and for the purposes of their report the procedure was deemed adequate.

To properly evaluate the effects of dam failure for the Midland project, this aspect of Brater and Wisler's work was revised. However, the basic rainfall - run-off results determined by Brater and Wisler were adopt-ed. The following paragraphs synopsize the pertinent aspects of Brater's and Wisler's work.

3.1 Probable Maximum Precipitation (PMP)

Based on USWB (1956), the PMP was determined to be 13.0 inches in 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months over 2,400 square miles preceded by 1.0 inch in the previous 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months occurring in June with no snowmelt.

The following table gives the probable maximum 24-hour

precipitations for eacn month. It also shows maximum potential snow-melts for 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months as determined by Brater and Wisler from their analysis of runoff and 3 temperature records in the Tittibawassee basin.

24-hour depth, inches 24-hour depth, inches PMP Melt Total PMP Melt Total JAN 4.3 1.0 5.3 JUL 12.9 -

12.9 l FEB 4.7 .8 5.5 AUG 12.8 -

12.8 l MAR 5.7 1.1 6.8 SEP 12.7 -

12.7 l APR 7.8 1.2 9.0 OCT 9.5 -

9.5 MAY 9.9 -

9.9 NOV 6.5 1.1 7.6 JUN 13.0 -

13.0 DEC 4.7 1.0 5.7 From the above, it is apparent that runoff from the June PMP would be greater than that from any storm accompanied by snowmelt.

3.2 Precipitation Losses In their derivation, Brater and Wisler determined in-filtration rates for each month based upon 25 storm runoff hydrographs. To incorporate the effects of antecedent moisture, intensity of precipitation, and the possible effects of ice on the ground, they used a lower envelope of abserved infiltration rates rather than the average observed values. The infiltration l thus derived varied from near zero for March to about l 0.8 inches per day for January. For June, the lower-

r bound filtration rate was found to be 0.4 inches per day. The initial abstraction was taken as 0.1-inches.

The resulting runoff from the 48-hour storm was thus 93 percent of the probable maximum precipitation.

3.3 Runoff Model Brater and Wisler developed unit hydrographs from the runoff and precipitation records of a number of large storms. These were done for 24-hour rainfalls and were very consistent between storms. Probable maximum rain-falls from USWB (1956) were applied to the resulting unit graph after modification for the losses discussed in section 3.2 above. The procedure was verified by application to three major floods at Midland. The re-sults were as follow:

3 Peak Discharge, cfs Flood Date Recorded Computed March 1942 26,100 26,000 March 1948 34,000 31,400 May 1912 48,000 52,000 l

4. NATURAL PROBABLE MAXIMUM FLOODING The hydrograph of the Tittabawassee River's natural PMF near the plant site was developed using the following, which were developed by Brater and Wisler (see Section 3 3, above):

1. The unitgraph shown in Figure 1

2. Probable maximum precipitation (PMP) and in-filtration losses as shown in Figure 2 Rainfall excesses were obtained by subtracting the loss-es from the rainfalls and were applied to the unitgraph to obtain the PMF hydrograph of Figure 2. This hydro-graph does not include the effects cf substantial over-bank storage which would occur just upstream of the 1 plant. (See Fig. 9.)

5. DAM FAILURE There are six existing hydroelectric power plant reser-voirs in the Saginaw River basin. All are located upstream of Midland, Michigan. Their characteristics 3 are as follows:

-M

EXISTING RESERVOIRS IN THE SAGINAW RIVER BASIN i \

Reservoir River Owner Drainage Gross Approx

** Area, Sq. Head Ft. Content Mile AF Secord T W 210 48 53,000 Smallwood T W 342 26 18,000 Edenville T W 985 43 84,000 Sanford T W 1020 28 32,000 Mt. Pleasant C H 325 13 ***

3 St. Louis P C 375 10 1,000

T = Tittabawassee, C = Chippewa, P = Pine

- W = Wolverine Power Co, H = Harris Milling Co.

C = City of St. Louis

= Not available, probably nil because the pond does not show on 1:250,000 topographic sheet.

(Since dam failure effects amounted to only about six percent of the PMF, failure of the dams independently of the PMF would result in levels lower than that cal-culated for the PMF.)

A reasonable mode and time of failure is required to evaluate the effect of dam failure upon the maximum 1 water level at the plant site. The worst condition would be if the four dams upstream from Sanford were to fail sucessively downstreamward. To allow for a possible " domino" effect in the failure of upstream 1 dams, the total storage behind all four dams could be treated as concentrated at the dam farthest downstream

-- Sanford Dam. Figure 3 gives the location and p'er-tinent features of Sanford Dam. 2,3 Study of the Sanford Dam design and records, and a brief site investigation indicated that failure would probably be by overtopping -- on which there is a dearth of usable information. A theoretical approach was tried for getting the logical mode and duration of failure by using the bed load transport ability of. flow-ing water. However, this did not give logical results after the first 20 to 30 minutes of the failure process.

A search of the technical literature of the past 15 years regarding dam failure by overtopping yielded only two instances where maximum discharge and failure duration 3 were given. The failure pattern finally selected for this study is shown in Figure 4. It assumes that:

4

. Failure start? at the low point near the riJ ht abutment

. The dam is completely removed in one hour

. The amount of material removed at intermediate 2 stages of degradation follows a geometric pro-gression. (This is reasonable because the rate af erosion increases with the area exposed to the flow, given that flow would be critical.)

It is logical that failure would start as the head-water level rises over the low point in the crest at the right end of the dam, and that the water level would continue to rise over the remaining longth of the crest due to two factors:

1. A high rate of inflow to the reservoir froa up-stream dam failure, in addition to the naturai flood. (The rate of inflow from upstream fail-ures was not studied since all storage to be re-leased was assumed concentrated at Sanford.]

2. Increasing restriction on the outflow through and over the dam -- from high tailwater levels that would result from a constriction in the channel about mile downstream at the C&O Rail- 2 road.

It was found that the latter condition would reduce the headwater-tailwater differential at the dam to less than 11 feet so that peak outflow through the break would be less than 210,000 cfs. It would also limit the amount I

of total stored water released through thc break to about 167,000 ac-ft. Figure 5 shows how this value was arrived at, and Figure 6 shows the hydrograph resulting from the dam failure. The rising limb of the failure hydrograph is a result of weir overflow calculations based on the assumed mode and duration of failure, and the recessive limb has been determined by trial, assuming all storage is released in about 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months .

The next step was to determine a reasonable point in time at which overtopping could be expected. Logically, this would occur when the rate of inflow to the reservoir exceeded the ultimate capacity of the spillway. This capacity was determined to be 25,000 cfs with headwater .5 feet over the low point in the dam crest. Sanford Dam spill-way has 6 gated bays, -2 @ 25.35' wide and 4 @ 22' wide.

Its ogee crest is 13.5' lower than the embankment low point and probably has a discharge coefficient of 4.0 at that head. Tailwater would be about 2.5' over the spillway crest and its effect is still negligible at that discharge.

The PMF hydrograph at Sanford Dam was then developed to obtain the time (from start of runoff) at which the discharge from Sanford would begin to exceed 25,000 cfs.

It was assumed outflow = inflow since the storage capability is small with respect to the flood volume. The hydrograph was developed using:

1. The unitgraph shown in Figure 7.

3

2. PMP and infiltration losses as shown in Figure 8.

Rainfall excesses were applied to the unitgraph to obtain the PMF hydrograph at Sanford Dam as shown in Figure 8. The hydrograph shows that overtopping would l3 i start about 1.5 days after the start of runoff.

Figure 9 (which is based on the USGS 15-minute topographic sheet "Sanford, Michigan") shows that a very extensive area just upstream of the plant would be inundated during an event euch as the PMF. Thus, average velocities would be quite low and storage effects quite high.

The failure hydrograph in Figure 6 was routed by the coefficient method thru the Tittabawassee River to the Midland plant site, using an average velocity of l 3.5 ft./sec., time increments of 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months , and channel 3 storage coefficient of .10. The assumed velocity and coefficient are considered appropriate for the PMF condition. !

l l

Figure 10 shows the channel-routed hydrograph along with the natural PMF at the plant site. The fact that the dam failure occurs early on the rising limb of the I flood, reduces the significance of the routing coefficients.

The dam breach is actually a relatively minor contribution to the total peak flow past the site.

l l

-C-

3 h

6. COMBINED PROBABLE MAXIMUM FLOOD As shown on Figure 10, adding the routed dam-failure hydrograph to the natural PMF at the site, produces a 3 combined peak flow of approximately 262,000 cfs - ,

8,000 cfs less than was previously estimated by Brater and Wisler (1956).

7. MAXIMUM WATER LEVEL It has been determined that the water level that would result at the plant site from the PMF peak of 262,000 cfs l in the Tittabawassee River would be about Elevation 630.1 - 3 say 631 in round numbers - under post-project conditions.

This is the result of rating curve calculations made using conservative assumptions regarding channel and flood plain flow resistance and downstream water levels.

Calculations were made using a USCE-originated computer program which uses the standard step backwater method.

For the calculations, the following were used:

1. Cross-sections obtained from four sources -

USGS 7.5 minute quadrangles with 5-foot contours, 1 foot contour maps by Abrams Aerial Survey, USCE surveys of 1948-49, and Bechtel surveys of 7.pril, 1974. Figures A-1 through A-5 of the appendix show these cross-sections.

2,3

2. Channel and flood plain roughness values at each cross-section as shown on Figure 11.

3. Starting water surface elevations for 5 different discharges arrived at by several trial runs for each discharge, selecting the higher of the pair of runs showing the best convergence.

For the cross-sections, the USGS maps were used generally above Elevation 610, sometimes higher. The use of USCE survey information was limited to the channel proper plus 200 to 300 feet of the flood plain. Abram's maps were used to fill in between USGS and USCE data, and Bechtel surveys were used for Sections 7 through 10.

In all, there were 11 cross-sections for computation of post-project water levels. The cross-section locations i are shown in Figure 11. I1' 3 For the purpose of developing judgement for "n" values in the area of concern, calculations were made to duplicate the water surface profile of the record flood of March, 1948 (34,000 cfs), checking at four points where the water level had been determined by the U.S. Corps of Engineers, Detroit District. The observed and the calculated water levels for pre-project conditions are listed in Table 1,

, together with Manning "n" values used in the calculations.

- ._ _ _ ~ _ . ___ _.. _. . . . _ _ . . _ . _ . , , .. . - _ . _.

._. . .-_ . __ _ - ~ . --_.- -- --

However, roughness values indicated by this study could not be applied blindly to the post-project conditions.

A study of recent aerial photographs showed th' for some reaches of the eastern flood plain it woulu be on the unsafe side to use the roughness indicated in the pre project study. In other reaches it would be ridicu-lously conservative to apply the "n" value determined in the pre-project study to the entire eastern flood plain.

The cross-sections were therefore divided into sub-sections and conservative but reasonable "n" values assigned to the eastern flood plain portions (Figure 11) . l3 The "n" value applied to the western flood plain is also conservative because construction of the cooling pond dikes eliminates the shallow, high "n" flood plain to the west as a flood flow area, leaving the deeper, lowe" "n" area to the east so that the right overbank "n" should

~

be lower than in any historical flood that inundated areas to the west. Also, because the computed pre-project water levels were slightly lower than the observed levels (Table 1), the channel "n" for cross-sections 10 through 6 was raised to 0.028 for post-project conditions.

To be conservative it was assumed that the new railroad

' bridge and the Smith's Crossing bridge (at Section 10 on l1,3 Figure 11) would not be destroyed by the PMF prior to the flood peak. However, Smith's Crossing bridge would almost certainly be destroyed. The result would be a lower-than-assumed water surface elevation at Section 10.

It was assumed that the railroad bridge-piers would gather debris so that the effective thickness of each pier would be about 3G feet.

Figure 12 shows the stage hydrograph which was obtained by interpolation among the various water levels 3 presented for "Section 4" in Table 1.

8. CONCURRENT WIND ACTIVITY l

It was postulated that an overland wind of 40 r.ph could i

occur from any direction during the PMF. The most adverse directions would.be from the northwest, over the " pend" shown on Figure 9, or from the southeast, directed upriver. 1 l3 In either case, the great extent of flooding would provide an effective fetch of 3.5 miles over which the wind would produce setup and waves.

Pertinent aspects of the two wind events are shown in Table 2.

I

\

,- The plant grade of 634 provides over 3 feet of freeboard against the still flood level, however, wind generated waves could carry water onto the site. For this reason, 3 an emergency procedure will be provided directing that critical openings in all Category I structures be sand-bagged under certain adverse conditions.

The following paragraphs discuss the development of the various items shown in Table 2. The principal reference 1'3 for this portion of the analysis is CERC (1973). It is referrred to henceforth as " Shore Protection Manual"(SPM) .

Maximum fetches (see Figure 9) and average depths were obtained from the USGS' 15-minute topographic sheets of 3 .

Sanford and St. Charles, Michigan and their 7.5-minute sheet Midland South, Michigan. Effective fetches were calculated using the sum-of-cesines approach shown on 1 3 Figure 3-14 of SPM. For ease of computation, 5-degree I segments were used over the 90-degree sector. (The construction for the upstream computation is shown on Figure 9 A similar one was employed downstream, orient-ing it to produce the maximum fetch.) l3 Overwater wind speeds were determined from Exhibit 12 of Reference 3 using the effective fetches. For the down- l3 stream wind effect, the stream velocity (as determined from earlier backwater computations) was added to the wind speed. But upstream, no stream effect was considered i

because of the great width of the pond.

Setups were determined from SPM's Equation 3-83 using the l3 effective wind, the maximum fetch, and the average depth.

To determine the featurcs of the significant waves, depths were obtained by considering the bottoms at Elevations 600 and 605 for the downstream and upstream events respectively.

From the topographic sheets, these elevations appear repre-sentative of the deeper regions where waves could form.

Setups were added to the PMF flood level of 631 before l3 obtaining depths.

Heights and periods of the significant waves were obtained 9 from SPM Equations 3-25 and 3-26 using these depths, effective fetch, and effective wind. Maximum wave heights were taken as 1.67 times the significant, in accordance with SPM Equation 7-2. Maximum wave periods were taken as 1.2 times the significant, based on Ippen (1966), 3 and minimum periods were determined from the breaking limit as presented in SPM Figure 7 41.

( For the wind from the north, potential runup was obtained as follows. With the full depth (to Elevation 605), and the deepwater wave lengths (5.12 times the square of the period), Table C-1 of SPM was used to get the equivalent deepwater wave height. With this, runup was determined 1 3 from SPM Figures 7-11 thru 7-13, assuming that smooth slopes of 3 horizontal to 1 vertical extend indefinitely upwards. The depth for determining runup is that between the setup stillwater level, 63L6 and the bottom, which was taken as 614. (A large area to the north and west of the plant fill is built up to this elevation.)

For the combined PMF/40-mph-wind event, some overtopping of the plant fill would occur over a significant portion 3 of the height-period spectrum. This is discussed in Section 9, below.

Haves coming from the south would be traveling upriver parallel to the affected plant slopes. Thus, there would be no runup in the conventional sense. The maximum height to which water would rise against these slopes was taken 1 as the elevation of the wave crests. These were obtained from SPM Figure 4-14. A search of the wave period spectrum indicated the highest crests would result from 3 a period of 2.8 seconds for the mnximum waves.

t Although these crests would exceed the plant grade, they would not produce significant quantities of overtopping flow, especially when compared with the massive, frontal attack produced by the wind blowing from the northwest. 3 Providing sufficient protection to keep the northerly waves from interfering with plant safety will obviously protect against the less severe southerly wave conditions.

t e

i 9.

FLOOD PROTECTION The final slope protection design will result from an y

1 economic evaluation of various alternatives. Current plans are to sod the embankments where they could possibly be exposed to wave action above the 100-year flood level.

Flooding protection itself will consist of sand-bagging the entrances of safety-related buildings to Elevation 635.5. To provide for timely sandbagging, the following guidelines should be followed in developing the flood protection emergency procedure.

O If the river reaches Elevation 625, mobilize the sandbagging effort (but do not necessarily put the sandbags.in place) and maintain an hourly watch on the river and the plant area.

o If the river reaches Elevation 627.8, maintain a continuous river watch.

o At the first sign of oversplash reaching any 3

-( . of the safety-related facilities, sandbag their entrances to Elevation 635.5.

An examination of Fig. 12 reveals that the rise in still 4

water level is sufficiently slow to allow ample preparation time. The elevation 625 allows approximately 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months to prepare for the placement of sandbags. The elevation 627.8 is the water level a* which the 40-mph wind's significant wave run up to Elevation 634. (All plant entrances are protected without sandbagging to Elevation 634.5.)

Sandbagging to Flevation 635.5 is based upon conservative approximations as to the behavior of water once it overtops the edge of the plant fill at Elevation 634. The greatest accumulation of overtopping flow would occur between the Warehouse and the Evaporator Building (see Fig. 13). These buildings will serve to funnel overtop across the railroad at that location. Because of the rapid increase in flow

(

t~

width downstream from this point, it was assumed that the railroad (at elevation 634.5) would serve as a weir. This would produce a backwater tcward the northwest which would restrict the quanity of overtopping flow.

The first calculation of this assumed a high weir coef"icient (3.4) and a high roughness factor (n) of .020. The backwater effect was calculated using standard step procedures and then solved together with the smooth-slope overtopping relationships presented by Wiegel (1964). The resulting flow past the buildings was found to be 190 cfs, which pro-duced a water surface near these two buildings at elevation 635.1.

3 The second calculation, assuming a low weir coeffecient of 1.7 and a roughness (n) of .012 (which is more indicative of predominantly asphalt area), resulted in a water surface at elevation 635.4 (flow = 165 cfs).

This indicates the relative insensitivity of the results to the data assumptions. It must be noted that these elevations are just upstream of the railroad. Substantial depth reductions will undoubtedly result from the great expansion once the water passes this point. However, for conservatisim, the safety-related buildings will be sand-( bagged to elevation 635.5 even though they are a considerable distance downstream of the point where these levels were calculated.

L

10. BULLOCK CREEK PMF Computations were made using 5 cross-sections and the standard step method to determine the maximum water level that would result if there were to be a Bullock Creek PMF concurrent with the arrival at the plant site of the 100-year flood peak on the Tittabawassee River. That maximum level was determined to be about Elevation 620 at the west dike of the cooling pond and poses no threat to vital installations since they are at Elevation 634. Figure A-6 of the Appendix shows l2 the channel cross-sections used. I The Bullock PMF was developed using:

1. The 2-hour unitgraph shown in Figure 14 l3

2. PMP and infiltration loss rate shown in Figure 15 Application of the maximum 24-hour rainfall excess to the unitgraph resulted in the PMF hydrograph in Figure 12.

The corresponding Bullock Creek water surface profile was calculated using the same computar program as used in the river backwater computations, with cross-sections

(- located as shown in Figure 16. l3 l

4

- 11. REFERENCES

1) Brater, E. F. and Wisler, C. '). " Maximum Flood Conditions at Midland, Michigan", McNamae, Porter & Seely Consulting Engineers; part of Dow Cher.ical Co. Report 56-21 Flood Protection of the MIdTand Plant from Extreme Floods on the Tittabawassee River, November, lW(unpublished) . --

2) Ippen, A. T., Estuary and Coastline Hydrodynamics, McGraw-Hill Book Co., Inc., New York, 1966.

3) OCE (Office of the Chief of Engineers), " Engineering and 3 Design -- Computation of Freeboard Allowances for Waves in Reservoirs", Engineer Technical Letter No. 1110-2-8, Rev. 1, Department of the Army, Washington, 16 December 1966.

4) USWB (U.S. Weather Bureau), " Seasonal Variation of Probable Maximum Precipitation East of the 105th Meridian....",

Hydrometeorological Report No. 33, Washington, April 1956.

5) Wiegel, R. L., Oceanographical Engineering, Prentice-! fall Inc., Englewood Cliffs, New Jersey, 1964.

(.

I

c ..

TABLE 1 OBSERVED AND CALCULATED WATER SURFACE ELEVATIONS SMITil'S CROSSING TO PLANT SITE AND USGS GAGE Section Station Pre-Project Post-Project No. W.S. Elevations Manning's "n" W. S. Elevations 34,000 cfs Obsrv. Calc. LOB CH RGB 55.000 75.000 100,000 170.000 270.000 10 0+00 606.59 606.59 0.150. 0.027 0.045 608.0 609.0 613.5 619.0 625.0 9 10F00 606.79 608.2 609.3 613.7 619.2 625.2 8 26+50 607.32 609.2 610.6 614.3 619.7 625.6 7 41+50 607.62 609.7 611.3 614.8 620.2 626.1 7.05 52+00 607.70 610.0 611.7 615.0 620.2 625.8 7.1 62+00 608.22 610.6 612.5 615.6 620.8 626.1 F' 6 76+00 608.56 608.39 0.040 0.027 0.045 611.0 613.0 616.2 621.7 627.7 Y 6.05 87+50 608.57 611.2 613.3' 616.6 622.4 628.6 6.07 '90+50 611.1 613.2 616.4 621.9 628.1 6.1 ds 93+50 611.1 613.1 616.3 621.7 628.0

. 6.1 us 94+00 611.4 613.7 619.8 623.2 628.0 5 102+50 603.51 611.6 614.0 620.3 624.1 629.1 5.1 121+00 609.03 612.1 614.6 620.7 624.8 630.0 2

42 144+50 609.10 609.15 0.040 0.027 0.06 612.2 614.8 620.9 625.1 630.4 23 179+50 609.78 609.63 Smith's Crossing Plant site -

USGS gage

TABLE 2 Wind Activity Concurrent with PMF Elevation of 631 t

Wind blowing from Upstream Downstream Maximum fetch, mi 7.9 211 Effective fetch, mi 3.5 ,

, 3.5 Overwater wind, mph 51 51 Stream velocity vector, nil +3 mph Effective wind, mph 51 54 Average depth, ft 20 28 Setup, ft 0.6 1.3 Depth for waves, ft 26.6 32.3 3

Wave heights, ft Significant 3.8 4.2 Maximum 6.3 6.9 Wave periods, sec Significant 3.8 4.0 Minimum 2.7 2.8 Maximum 4.0 4.8 Potential Runup Elevation (See Section 8) at significant wave 637.9 634.6 at maximum wave, worst period 641.6 637.0

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