Attach_5_Supplemental_Info-Tech_Report-Scorpion_Pier_Coastal_Engr_Analysis__03-15-11.pdf

Technical Report

COASTAL ENGINEERING ANALYSIS

SCORPION ANCHORAGE PIER REPLACEMENT PROJECT

Prepared for U. S. National Park Service

March 15, 2011

Version Preparation QC Review

Draft 1 David Simpson Scott Fenical

Final David Simpson Scott Fenical

Coastal Engineering Analysis - Scorpion Anchorage Pier Replacement Project March 15, 2011

TABLE OF CONTENTS

1. Introduction

2. Pier Alternatives

3. Project Site Physical Conditions

3.1. Tides

3.2. Sea Level Rise

3.3. Winds

3.4. Waves

3.4.1. Summary of Buoy Data

3.4.2. Wave Model Setup and Validation

3.4.3. Extreme Waves

4. Wave Loading on Marine Structures

5. Wave Breaking Location

6. Coastal Flooding Analysis

7. Shore Protection Design Criteria

8. Conclusions

9. References

Appendix A – Inundation Levels for Sea Level Rise Scenarios

FIGURES

Figure 1. Locations of replacement pier alternatives (from Jones & Jones, Inc.)

Figure 2. Measured NOAA sea level data and projections by NOAA and California Climate

Action Team (CCAT, 2010)

Figure 3. Wind speed and direction distribution (wind rose) at NDBC Buoy 46053

Figure 4. Extreme sustained 2-minute wind speeds and directions for return periods between 1 and 100 years

Figure 5. National Data Buoy Center (NDBC) buoys in Santa Barbara Channel

Figure 6. SWAN large-area bathymetry domain and location of nested domain shown in Figure

3-4

Figure 7. SWAN nested modeling domain

Figure 8. SWAN model validation results for largest Pacific Ocean storm recorded at Buoy

46218, February 24, 2008

Figure 9. SWAN model significant wave heights at Scorpion Anchorage during Pacific Ocean storm recorded on February 24, 2008

Figure 10. SWAN model significant wave heights at Scorpion Anchorage during largest east wind event on record at NDBC Buoy 46053 (February 11, 2003)

Figure 11. Horizontal wave force as a function of circular pile diameter

Figure 12. Vertical uplift force as a function of deck elevation

Coastal Engineering Analysis - Scorpion Anchorage Pier Replacement Project March 15, 2011

Figure 13. Inundation by MHHW tide levels for current sea level and projected relative increases for years 2050 and 2100

Figure 14. Definition sketch of wave runup and overtopping at a low bluff (source: Cox and

Machemehl (1987))

Figure 15. Coastal flooding analysis transect locations

Figure 16. Wave runup elevation and extent of inundation by 100-year swell wave overtopping occurring at times of MHHW for years 2010, 2050, and 2100

Figure 17. Wave-induced inundation extents during 100-year storm (swell) at MHHW tidal elevation in 2010 (blue), in 2050 (green), and in 2100 (red)

TABLES

Table 1. Tidal datums for Becher's Bay, Santa Rosa Island

Table 2. Measured rates of sea level rise in vicinity of Santa Cruz Island

Table 3. Wave height, period, and direction at pier locations for extreme Pacific Ocean storms.11

Table 4. Wave height, period, and direction at pier locations for extreme easterly storms

Table 5. Wave breaking locations and maximum crest elevations at MLLW tide level for extreme Pacific Ocean storms

Table 6. Wave breaking locations and maximum crest elevations at MLLW tide level for extreme easterly wind-waves

Table 7. Wave breaking locations and maximum crest elevations at MHHW tide level for extreme Pacific Ocean storms

Table 8. Wave breaking locations and maximum crest elevations at MHHW tide level for extreme easterly wind-waves

Table 9. Wave runup elevation and upland inundation extents by wave overtopping

Table 10. Stable armor stone specifications for 100-year design storm with various revetment design slopes

Technical Report Page 1 Coastal Engineering Analysis – Scorpion Anchorage Pier Replacement Project March 15, 2011

Coastal Engineering Analysis Scorpion Anchorage Pier Replacement Project

1. Introduction

Coast & Harbor Engineering, Inc. (CHE) performed coastal engineering analysis for the

Scorpion Anchorage Pier Replacement Project for the U. S. National Park Service as part of a team with Jones & Jones, Inc. (Project Team). The project site is located at Scorpion

Anchorage on Santa Cruz Island, as shown in Figure 1. The project goal is to enhance safety for park visitors and reduce facility maintenance at the project site. CHE performed the coastal engineering analysis to provide design criteria and assist the Project Team in development of design concepts.

The coastal engineering analysis consisted of the following elements:

� Bathymetry/topography data compilation and analysis

� Coastal conditions evaluation (tides, sea level rise, winds, waves)

� Wave loading on marine structures (piles, deck)

� Coastal flooding evaluation (wave runup and overtopping)

� Shore protection design criteria (stable rock size)

2. Pier Alternatives

Three replacement pier conceptual alternatives were provided to CHE at the time of this report. Figure 1 shows the proposed replacement pier alternative locations. The replacement pier alternatives were assumed to be pile-supported. Design wave conditions and loads on marine structures were developed for the seaward end of the pier at each alternative location.

3. Project Site Physical Conditions

3.1. Tides

Tides at the Scorpion Anchorage are mixed semi-diurnal. Tide data of locations surrounding the site were collected from Prisoner's Harbor (NOAA Station ID #467), Santa Barbara (NOAA Station ID #9411340), and Becher's Bay, Santa Rosa Island

(NOAA Station ID #9410962). Tides at Becher's Bay, Santa Rosa Island were determined to be the most representative of the site, and tidal datums are shown in

Table 1.

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Figure 1. Locations of replacement pier alternatives (from Jones & Jones, Inc.)

Table 1. Tidal datums for Becher's Bay, Santa Rosa Island

Datum Elevation (ft, MLLW)

MHHW 5.14

MHW 4.41

DTL 2.57

MTL 2.69

MSL 2.67

MLW 0.96

MLLW 0.00

3.2. Sea Level Rise

Sea level at any location depends on numerous long-term and short-term processes acting at regional and global scales. Relative sea level rise at the project site is the combination of vertical movement of the local land mass, and eustatic (global) changes in sea level. Vertical land movement (rise) at Santa Cruz Island is estimated to be +0.7 to +1.5 mm/yr (Selander et al., 2011).

Short-term factors that affect water levels at the project site include seasonal fluctuations, El Niño or La Niña effects, barometric pressure gradients, and waves.

The time scale of these factors is typically days to months. The focus of the present

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SLR analysis is long-term trends. Historical local rates of sea level rise (SLR) were estimated from sea level measurements at nearby NOAA tide stations. NOAA publishes historical rates of sea level change at five nearby stations along the coast in the region around Santa Cruz Island (see Table 2), with long-term SLR estimates ranging from +0.79 to +3.22 mm/year.

Table 2. Measured rates of sea level rise in vicinity of Santa Cruz Island

Location Sea Level Rise (mm/year)

Port San Luis 0.79

Santa Barbara 1.25

Rincon Island 3.22

Santa Monica 1.46

Los Angeles 0.83

Source: http://co-ops.nos.noaa.gov/sltrends/index.shtml

Measured SLR at the Rincon Island Station (NOAA Station ID 9411270) shows a trend of 3.22 +/- 0.99 mm/year based on 28.7 years of data between 1962 and 1990

(Pendleton et al., 2005).

Predictions of SLR due to climate change are based on results of computer modeling with some of these results reported in CCAT (2010). SLR predictions based on these climate change models vary over a wide range. Accelerated SLR, predicted to occur due to climate change or melting of land-locked ice sheets, such as those in

Antarctica or Greenland, is not noticeable in the measured water level records at these stations. Figure 2 shows the range of CCAT projections (Low, Medium and High) and the historical SLR trend at Rincon Island with base year 2000. Because the

CCAT projections of SLR are higher than the observed SLR even by the time of this report, the CCAT projections should be used only with an understanding of the likely conservatism involved.

To provide a reasonably conservative estimate of relative SLR, we assume the low range of projected vertical land mass rise at Santa Cruz Island (+0.7 mm/year) and examine the range of observed and projected SLR. Relative SLR at the project site based on historical trends is extrapolated to be 0.33 feet by 2050 and 0.74 feet by

2100. Relative SLR at the project site based on "High" CCAT projections, as an example of a high projection, is 1.1 feet by 2050 and 4.5 feet by 2100.

The "Low" CCAT projection is recommended for the purpose of evaluating coastal flooding at the project because it represents a mean between the highest rate of accelerated sea level rise and the observed constant rate. The relative SLR at the project site is 1.1 feet by 2050 and 3.2 feet by 2100 based on "Low" CCAT projection and the low estimate of rate of rise of the land.

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0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

5.0

2000 2010 2020 2030 2040 2050 2060 2070 2080 2090 2100

Year

E u s ta ti c S e a L e v e l R is e fe e t)

CCAT High

CCAT Med

CCAT Low

NOAA Rincon Island

Figure 2. Measured NOAA sea level data and projections by NOAA and California Climate Action Team (CCAT, 2010)

3.3. Winds

An extreme value analysis of winds near the project site was performed for developing coastal engineering design criteria. Wind data from local buoys were obtained and processed, including NDBC Buoy 46053 (119.841000° W, 34.248000° N) located north of Santa Cruz Island in Santa Barbara Channel. Good data quality and the completeness and length of record (1995 through 2009) indicated that the NDBC Station 46053 dataset should be used. Figure 3 shows the distribution of measured wind speed and direction at NDBC Station 46053 in the form of a wind rose, with winds corrected to 10 m elevation and 2-minute-average speed. The wind rose shows the predominance of winds from the west at the buoy location.

The largest measured wind speeds were extracted from the record for all directions at

10-degree increments. These extreme speeds were fit to a Weibull distribution, and sustained wind speeds were predicted for extreme events with return periods between

1 and 100 years for all directions. Figure 4 shows the predicted extreme wind speeds for given return periods and directions.

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Figure 3. Wind speed and direction distribution (wind rose) at NDBC Buoy 46053

Figure 4. Extreme sustained 2-minute wind speeds and directions for return periods between 1 and 100 years

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3.4. Waves

3.4.1. Summary of Buoy Data

The wave environment at Scorpion Anchorage is energetic due to its exposure to waves generated in the Santa Barbara Channel by strong winds and to Pacific Ocean waves arriving from deep water and refracting around the northeast end of the island.

Wave measurements are available from several locations in the vicinity of the Santa

Barbara Channel. Figure 5 shows the locations of National Data Buoy Center

(NDBC) buoys in the area. Wave records were compiled, analyzed and quality-checked to ensure accurate extreme wave analysis. Wave records from

NDBC 46218 (Harvest Buoy) contained the longest period of record (approximately

15 years) and best represented incoming Pacific Ocean waves, and therefore were used for all wave transformation modeling and analysis. A record length of at least

30 years is typically desired for predicting the 100-year recurrence interval; however, no datasets of that length are available.

Figure 5. National Data Buoy Center (NDBC) buoys in Santa Barbara Channel

3.4.2. Wave Model Setup and Validation

Wave transformation analysis was performed with the two-dimensional (2-D) spectral model SWAN (Holthuijsen et al., 2004) to determine extreme wave parameters at the project site. SWAN is a phase-averaged spectral wave transformation model that is well suited for use on this project due to its capability to incorporate nested modeling domains, allowing accurate propagation from large-area to small-area domains. The model includes generation of wave energy due to wind, wave transformation, wave-wave interaction, and energy dissipation due to breaking and white capping.

Modeling was performed in steady-state mode with a spatially and temporally constant input wind field, and spatially constant input wave field with the offshore

Technical Report Page 7 boundary located near NDBC Buoy 46218. The results of the SWAN model included spatial representations of significant wave height (average of the highest one-third of the waves) and wave spectral peak period, as well as other parameters.

Figure 6 shows depths in the large-area SWAN modeling domain, constructed using the Santa Barbara tsunami modeling digital elevation model (DEM) produced by

NOAA (Carignan et al., 2009). Figure 7 shows the SWAN nested grid, constructed with the NOAA tsunami DEM as well as 2006 multi-beam data (California State

University Monterey Bay Seafloor Mapping Lab) and nearshore single-beam survey data collected in October 2010 by Merkel & Associates.

Figure 6. SWAN large-area bathymetry domain and location of nested domain

Figure 7. SWAN nested modeling domain

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For SWAN model validation, wave data from NDBC 46218 were reviewed and the largest westerly storm conditions were extracted. The largest west waves on record occurred on February 24, 2008 and consisted of significant wave height of 32.7 feet, and peak period 18.2 seconds arriving from 280 degrees (measured from True North).

Local peak wind speed at the time of the peak swell was 10.4 knots from 250 degrees

(True North).

Wind and wave time histories recorded at the buoy were input to the SWAN model to simulate a 36-hour period starting February 24, 2008 00:00 (UTC). Results of the model validation are shown in Figure 8 in the form of significant wave height time histories at the measurement buoy (46218), with four other buoys in the Santa

Barbara Channel. Buoy data indicate that wave dissipation occurs as Pacific Ocean swell propagates into the Channel, and this dissipation is well represented by the wave transformation model.

3.4.3. Extreme Waves

Extreme waves at the project site arrive from two sources: (1) Pacific Ocean swell and westerly winds, and (2) north/northeast winds (Santa Ana winds). These wind and wave systems cause different impacts at the existing pier due to the different wave characteristics. Pacific Ocean waves are longer in period, and hence, are likely to cause larger wave runup and coastal flooding. Local wind-waves are known to cause berthing difficulties at the existing pier, and may control wave loads on marine structures. Therefore, extreme waves from both wave systems were modeled and are reported here.

In general, the two wave systems are independent. Pacific Ocean swell are typically generated far from the area, and may or may not arrive at the same time as easterly wind-waves. Combining extreme swell conditions and extreme easterly wind-wave conditions results in joint occurrence return periods that are much larger than that recommended for pier design.

To account for the complexity of the two wave systems and their joint probability, a response-based approach was employed to determine extreme wave conditions at the terminal alternative locations. The response-based approach consists of simulating a large number of actual historical combinations of swell and winds with their historically accurate tidal elevation. This procedure results in a large set of extreme wave conditions at the pier which can then be input to an extreme value analysis procedure to determine return period wave heights.

To develop this large database of actual storm events, the 100 largest Pacific Ocean wave events were extracted from the buoy records, along with their concurrent west winds and tidal elevations. Independently, the 100 largest easterly winds were extracted from the buoy record, along with their concurrent Pacific Ocean swell and tidal elevations. All 200 of these actual storm conditions were simulated using the

SWAN model with predicted tidal elevations.

Figure 9 shows example SWAN modeling results near Scorpion Anchorage for the largest Pacific Ocean storm on record for buoy 46218 (February 24, 2008).

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Significant wave height at the existing pier is approximately 3.8 feet with a peak period of 19.0 seconds.

Figure 10 shows example SWAN modeling results near Scorpion Anchorage for the largest east wind event on record at NDBC Buoy 46053 (February 11, 2003).

Significant wave height at the existing pier is approximately 7.2 feet with a peak period of 8.2 seconds.

Figure 8. SWAN model validation results for largest Pacific Ocean storm recorded at Buoy 46218, February 24, 2008

2/24

12:00

2/24

18:00

2/25

0:00

2/25

6:00

2/25

12:00

2/25

18:00

2/26

0:00

Date & Time (UTC)

Significant Wave Height (feet)

Buoy Locations

46218

46054

46053

46216

46217

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Figure 9. SWAN model significant wave heights at Scorpion Anchorage during Pacific Ocean storm recorded on February 24, Figure 10. SWAN model significant wave heights at Scorpion Anchorage during largest east wind event on record at NDBC Buoy 46053 (February 11, 2003)

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After the most severe 200 historical wind, wave, and tide conditions were simulated with the SWAN model, the resulting significant wave heights and peak periods were extracted at the seaward ends of each pier alternative. The 30 largest swell-based wave conditions (from the west) and the 30 largest wind-wave based wave conditions

(from the east) in terms of significant wave height were extracted from the storm modeling results. The 30 events were fit to a Weibull distribution, and extreme events between 2 and 100 years were predicted. A one-year return period value was obtained by extrapolation.

Because the top 30 westerly storms (swell) and top 30 easterly storms (wind-waves) were extracted separately, extreme wave heights and periods can be reported for swell and for wind-waves independently. Care was taken to identify extreme waves that were dominated by swell and separately those dominated by east winds because they need to be analyzed separately. This was performed because swell waves control runup and overtopping, while larger wind-waves control loads on structures. Tables 3 and 4 list predicted extreme height and period for swell and wind-waves, respectively.

The representative wave period applied for all extreme swell events (19.0 seconds) was estimated based on review of measured storm conditions. Wave periods for wind-wave events were developed through additional SWAN modeling runs, in which modeled wave heights were matched with extreme predictions of wave height.

Table 3. Wave height, period, and direction at pier locations for extreme Pacific Ocean storms

Alternative 1 Alternative 2 Alternative 3 Return Period (years) Hs (ft) Tp (s) θ (deg)

Hs (ft) Tp (s) θ

(deg) Hs (ft) Tp (s) θ (deg)

1 3.3 19 25 3.0 19 25 3.0 19 25

2 3.5 19 25 3.2 19 25 3.2 19 25

5 3.7 19 25 3.4 19 25 3.5 19 25

10 3.9 19 25 3.6 19 25 3.6 19 25

25 4.1 19 25 3.8 19 25 3.9 19 25

50 4.2 19 25 3.9 19 25 4.0 19 25

100 4.4 19 25 4.0 19 25 4.1 19 25

Note: θ is local direction from which waves approach, measured clockwise from True North, in degrees.

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Table 4. Wave height, period, and direction at pier locations for extreme easterly storms

Alternative 1 Alternative 2 Alternative 3 Return Period (years) Hs (ft) Tp (s) θ (deg)

Hs (ft) Tp (s) θ

(deg) Hs (ft) Tp (s) θ (deg)

1 5.8 6.2 80 5.5 6.2 70 5.2 6.2 65

2 6.3 6.6 80 5.9 6.6 70 5.7 6.6 65

5 6.9 7.0 80 6.5 7.0 70 6.2 7.0 65

10 7.3 7.3 80 6.8 7.3 70 6.6 7.3 65

25 7.8 7.7 80 7.2 7.7 70 7.0 7.7 65

50 8.2 8.0 80 7.6 8.0 70 7.3 8.0 65

100 8.6 8.2 80 7.9 8.2 70 7.6 8.2 65

Note: θ is local direction from which waves approach, measured clockwise from True North, in degrees.

4. Wave Loading on Marine Structures

Forces on pier structures were calculated for the largest wave in a storm having a 100-year return period occurring at a time of MHHW tide level. Wave-induced forces on the pier structures were generated using the results of wave transformation modeling and empirical loading methods. Horizontal wave loads on the vertical support piles (assumed to be circular) were generated using significant wave heights calculated from SWAN results in

Section 3 (converted to maximum wave heights assuming a Rayleigh distribution). Wave kinematics for these wave conditions were calculated using Stream Function wave theory

(Dean 1965), and horizontal loads on circular piles were calculated using the formulation of

Morison et al. (1950).

Vertical wave forces (uplift) on the pier deck were evaluated using the same input wave and hydrodynamic information as input to the formulation of Cuomo et al. (2007). Extreme waves generated by east winds are the design waves for loading on pier structures. Figure 11 provides the peak horizontal wave loads on vertical circular piles as a function of diameter for the three different pier alternative locations. Figure 12 is a plot of peak vertical wave uplift forces on the horizontal decks as a function of deck elevation for the three different pier alternative locations. Figure 12 provides values for quasi-static force and for impact force. Quasi-static force is uplift occurring over a wider area of the deck and should be applied in a temporally constant manner for design of large pier components. Impact force is a slamming-type of uplift which occurs over very small areas.

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1,000

2,000

3,000

4,000

5,000

6,000

7,000

8,000

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Circular Pile Diameter (feet)

T o ta l F o rc e lb f)

Alternative 1

Alternative 2

Alternative 3

Figure 11. Horizontal wave force as a function of circular pile diameter

0 250 500 750 1,000 1,250 1,500 1,750 2,000

Uplift Force on Deck (lbf/ft

D e c k B o tt o m E le v a ti o n

(f e e t, M L L W

Alt 1 Quasi-Static Force

Alt 2 Quasi-Static Force

Alt 3 Quasi-Static Force

Alt 1 Impact Force

Alt 2 Impact Force

Alt 3 Impact Force

Figure 12. Vertical uplift force as a function of deck elevation

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5. Wave Breaking Location

Wave breaking locations were determined since they may be useful for anticipating potential impacts to surface-piercing structures and moored vessels. Wave breaking locations were calculated in terms of bottom elevation for maximum wave heights in each storm spectrum for each return period. Separate calculations were made for tide levels at MLLW and

MHHW. Wave breaking is possible at locations between the selected tide levels. Bottom elevations at which wave breaking begins at MLLW tide level are shown in Tables 5 and 6 for extreme Pacific Swell and extreme easterly wind storms, respectively. Similarly, Tables 7 and 8 list bottom elevations corresponding to wave breaking at MHHW tide level.

Table 5. Wave breaking locations and maximum crest elevations at MLLW tide level for extreme Pacific Ocean storms

Alternative 1 Alternative 2 Alternative 3

Return Period (years)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest Elevation (ft, MLLW)

1 -7.0 3.8 -6.5 3.5 -6.5 3.5

2 -7.5 4.1 -6.9 3.7 -6.9 3.7

5 -8.0 4.3 -7.3 3.9 -7.5 4.1

10 -8.4 4.5 -7.7 4.2 -7.7 4.2

25 -8.8 4.8 -8.2 4.4 -8.4 4.5

50 -9.0 4.9 -8.4 4.5 -8.5 4.7

100 -9.4 5.2 -8.5 4.7 -8.8 4.8

Table 6. Wave breaking locations and maximum crest elevations at MLLW tide level for extreme easterly wind-waves

Alternative 1 Alternative 2 Alternative 3

Return Period (years)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest

Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest Elevation (ft, MLLW)

1 -14.5 8.3 -13.5 7.8 -12.8 7.3

2 -15.6 8.9 -14.4 8.3 -13.8 8.0

5 -16.9 9.7 -15.8 9.1 -14.9 8.6

10 -17.9 10.3 -16.4 9.5 -15.9 9.1

25 -18.9 10.9 -17.3 10.0 -16.7 9.7

50 -19.9 11.4 -18.1 10.5 -17.4 10.0

100 -20.8 12.0 -18.9 10.9 -18.0 10.4

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Table 7. Wave breaking locations and maximum crest elevations at MHHW tide level for extreme Pacific Ocean storms

Alternative 1 Alternative 2 Alternative 3

Return Period (years)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest

Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest

Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest Elevation (ft, MLLW)

1 -1.9 8.9 -1.4 8.6 -1.4 8.6

2 -2.4 9.2 -1.8 8.8 -1.8 8.8

5 -2.9 9.4 -2.2 9.0 -2.4 9.2

10 -3.3 9.6 -2.6 9.3 -2.6 9.3

25 -3.7 9.9 -3.1 9.5 -3.3 9.6

50 -3.9 10.0 -3.3 9.6 -3.4 9.8

100 -4.3 10.3 -3.4 9.8 -3.7 9.9

Table 8. Wave breaking locations and maximum crest elevations at MHHW tide level for extreme easterly wind-waves

Alternative 1 Alternative 2 Alternative 3

Return Period (years)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest

Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest

Elevation (ft, MLLW)

Wave Breaking Bottom Elevation (ft, MLLW)

Wave Crest

Elevation (ft, MLLW)

1 -9.4 13.4 -8.4 12.9 -7.7 12.4

2 -10.5 14.0 -9.3 13.4 -8.7 13.1

5 -11.8 14.8 -10.7 14.2 -9.8 13.7

10 -12.8 15.4 -11.3 14.6 -10.8 14.3

25 -13.8 16.0 -12.2 15.1 -11.6 14.8

50 -14.8 16.5 -13.0 15.6 -12.3 15.1

100 -15.7 17.1 -13.8 16.0 -12.9 15.5

6. Coastal Flooding Analysis

Inundation of the project area by future sea level increases (without wave action) is illustrated in Figure 13. The figure shows the MHHW contour for existing sea level and for cases of the 1.1-ft relative rise projected for 2050 and the 3.2-ft relative rise projected for

2100, assuming topography remains similar to the current configuration. Individual figures showing the MHHW inundation levels are in Appendix A.

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Figure 13. Inundation by MHHW tide levels for current sea level and projected relative increases for years 2050 and 2100

Coastal flooding analysis was performed for extreme waves using the results of the SWAN modeling and empirical methods of calculating wave runup and overtopping of the upland crest presented by FEMA (2005). Figure 14 illustrates the overtopping process and defines terms of wave runup elevation and inundation zone. The highest waves in the storm associated with the coastal flood were assumed to coincide with a tidal elevation of MHHW

(5.14 ft above MLLW). Wave runup elevation was characterized by the two percent highest runup height during the time of maximum waves. Comparison of runup values from the two sources of waves showed that swell waves control coastal flooding.

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Pacific swell corresponding to return periods of 25, 50, and 100 years (Table 3, Alternative 1) were input to the runup and overtopping analysis procedure. Beach profiles were generated at four representative locations along the beach, as shown in Figure 15, and beach profile parameters were also input to the procedure. Coastal flooding analysis determined the inland distance along the four beach profiles reached by the overtopping wave during the specified wave and tide conditions. This distance defines the inundation zone, whether caused by runup on a coastal bluff or by overtopping of a beach berm or low upland. Wave overtopping by refracted Pacific swell waves having a 100-year return period was compared with overtopping due to 100-year east wind-wave storms. Results showed that flooding by wave overtopping was controlled by swell runup.

Figure 14. Definition sketch of wave runup and overtopping at a low bluff (source:

Cox and Machemehl (1987))

Figure 15. Coastal flooding analysis transect locations

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Figure 16 illustrates the cross-shore profiles at the four transect locations shown in Figure 15.

Figure 16 also shows the MHHW elevation at present and projected elevations in 2050 and

2100, and the 100-year runup elevations and inundation zones calculated for the three mean water levels. Runup elevations apply only to locations near the shore. As illustrated in

Figure 14, the wave that runs up the beach and overtops the upland crest diminishes in height as it runs over land.

Runup elevations relative to MLLW and inundation zone widths measured from the MHHW contour line established by the 2010 survey are listed in Table 5 for the four transects and three sea level scenarios for extreme storms. Figure 17 shows the areas of maximum inundation of the upland by wave action (coastal flooding) during the 100-year storm at the current sea level, in 2050 after 1.1 ft of relative sea level rise, and in 2100 after 3.2 ft of relative sea level rise. Maps of areas impacted by seawater inundation by coastal flooding are shown in Appendix A for the 100-year storm with associated waves and wave setup, each for given sea level increases of 1.1 ft and 3.2 ft.

a) Transect 1 b) Transect 2

c) Transect 3 d) Transect 4

Figure 16. Wave runup elevation and extent of inundation by 100-year swell wave overtopping occurring at times of MHHW for years 2010, 2050, and 2100

Runup elevations and overtopping distances do not vary greatly at a transect for the range of return periods at the Scorpion Anchorage, as observed in Table 9, for two reasons. The first reason is the large amount of diffraction that all the swell waves undergo in reaching the

Technical Report Page 19 location of the end of the pier. Diffraction of these waves produces nearly the same wave height there. The second reason is the long wave length of the extreme swell dominates the overtopping process. The similarity of wave period in these extreme swell events causes the inundation limit to also be similar for the range of return periods and a given sea level.

Therefore, for each sea level scenario the inundation limits of the 25-, 50-, and 100-year wave storms are so nearly the same that the plot of only the 100-year inundation is shown in

Figure 17 and in Appendix A.

The large inundation extent shown for Transects 2 and 3 for the year 2100 results from the assumed sea level rise. The projected MHHW elevation in 2100 will lead to overtopping of the beach crest and flooding of the low area landward of the beach berm. This is likely a conservative scenario because the berm crest is formed and maintained by dynamic coastal processes. The crest elevation is expected to rise as average runup elevation increases through time. It is understood that the creek floods through the same area regularly and is probably the dominant flooding source in the low upland area.

Table 9. Wave runup elevation and upland inundation extents by wave overtopping

Transect Current Sea Level 1.1 ft Sea Level Rise 3.2 ft Sea Level Rise

25-Year Wave Runup Elevation (ft) / Overtopping Distance (ft)

Runup Overtopping Runup Overtopping Runup Overtopping

1 11.50 27 12.60 38 14.70 46

2 10.42 73 11.52 93 13.62 307

3 10.40 39 11.50 60 13.60 346

4 10.43 48 11.53 54 13.63 63

50-Year Wave Runup Elevation (ft) / Overtopping Distance (ft)

Runup Overtopping Runup Overtopping Runup Overtopping

1 11.59 27 12.69 39 14.79 46

2 10.49 73 11.59 93 13.69 307

3 10.47 41 11.57 60 13.67 346

4 10.51 49 11.61 55 13.71 63

100-Year Wave Runup Elevation (ft) / Overtopping Distance (ft)

Runup Overtopping Runup Overtopping Runup Overtopping

1 11.65 27 12.75 39 14.85 46

2 10.54 73 11.64 93 13.74 307

3 10.52 46 11.62 60 13.72 346

4 10.56 49 11.66 55 13.76 63

Notes: Elevation referenced to MLLW of current epoch Coastal wave flooding calculated with tide level of MHHW for corresponding year

Extents measured as horizontal distance landward of MHHW contour (2010 survey)

Technical Report Page 20

Figure 17. Wave-induced inundation extents during 100-year storm (swell) at MHHW tidal elevation in 2010 (blue), in 2050 (green), and in 2100 (red)

7. Shore Protection Design Criteria

Rock shore protection design criteria were developed for the 100-year storm. Waves generated by winds blowing from the east (rather than Pacific swell) control shore protection design. A design significant wave height of 8.53 ft and period of 8.2 seconds were input to the armor stone stability formulation of van der Meer (1988) to develop recommended rock sizes and weights. Table 10 shows the recommended armor stone size and weight for the

100-year return period storm for rock revetment slopes between 1.5H:1V and 3H:1V.

Technical Report Page 21

Computed stone sizes compare well with the size of naturally-occurring rock at the shore adjacent to the existing pier. Only median stone size is provided in Table 10. Shore protection design will require further specification of armor and bedding stone weight distributions (gradations), rock quality, and other design components.

Table 10. Stable armor stone specifications for 100-year design storm with various revetment design slopes

Slope (H:V) D50 (ft) W50 (lbf)

1.5:1 4.0 6,030

2:1 3.7 4,830

3:1 3.0 2,630

8. Conclusions

Coastal engineering analysis and numerical modeling were performed as part of the Scorpion

Anchorage Pier Replacement Project. The objectives of these analyses were to develop coastal design criteria for pier alternatives and background data for future shore protection design.

Site conditions and the coastal engineering analysis indicate that the design wave climate at the project site is energetic and likely to have an impact on replacement pier design. Design criteria provided here should provide a basis for marine structure analysis and future refinement of the location and design of the replacement pier.

Coastal engineering parameters presented in this Technical Memorandum are design wave forces corresponding to given circular pile diameters and to pier deck elevation, coastal flooding zone width for current sea level and two assumed cases of sea level rise, and median size of rock for use in erosion control at the shoreline for a range of revetment slopes.

9. References

Carignan, K. S., L. A. Taylor, B. W. Eakins, R. R. Warnkin, E. Lim, and P. R. Medley. 2009.

Digital Elevation Model of Santa Barbara, California: Procedures, Data Sources and

Analysis. NOAA Technical Memorandum NESDIS NGDC-29, Boulder CO.

CCAT. October 2010. State of California Sea-Level Rise Interim Guidance Document.

Sea-Level Rise Task Force. Coastal and Ocean Working Group of the California Climate

Action Team. California Environmental Protection Agency. Sacramento, CA.

Cox, J. C. and J. Machemehl. 1987. Overland Bore Propagation Due to an Overtopping Wave.

Journal of the Waterways, Ports, Coastal and Ocean Engineering Division, American

Society of Civil Engineers, Vol. 112, No. 1, p. 161 – 163.

Cuomo G., Tirindelli M. & Allsop N.W.H. 2007. Wave-in-Deck Loads on Exposed Jetties.

Coastal Engineering, Vol. 54 (9), 657-679.

Technical Report Page 22

Dean, Robert G. 1965. Stream Function Wave Theory: Validity and Application. Proceedings of Coastal Engineering, Specialty Conference of American Society of Civil Engineers, Santa Barbara, CA, October 1965, p. 269-300, A.S.C.E., New York

FEMA. 2005. Guidelines for Coastal Flood Hazard Analysis and Mapping of the Pacific Coast of the United States. Prepared for FEMA by Northwest Hydraulic Consultants. FEMA

Region 9, Oakland, CA.

Holthuijsen, L.H., Booij, N., Ris, R.C., Haagsma, IJ.G., Kieftenburg, A.T.M.M., Kriezi, E.E., Zijlema, M. and van der Westhuysen, A.J. February 5, 2004. SWAN Cycle III Version

40.31 User Manual.

Intergovernmental Panel on Climate Change (IPCC). 2007. Global Climate Predictions

(Chapter 10). The Intergovernmental Panel on Climate Change, United Nations, New

York, NY.

Morison J.R, O’Brien M.P.O., Johnsen J.W. and Schaff S.A. 1950. The forces exerted by surface waves on piles. Petrol Trans. AIME, vol. 189, pp. 149-157.

Pendleton, Elizabeth, E. Thieler, and S. Williams. 2005. Coastal Vulnerability Assessment of

Channel Islands National Park (CHIS) to Sea-Level Rise (Open-File Report 2005-1057).

U.S. Geological Survey. Reston, Virginia.

Selander, Jacob and A. Clark. Geology and Geomorphology of Eastern Santa Cruz Island Field

Guide-Chapter 4 Neotectonics of Santa Cruz Island, California. Department of Geology.

University of California. Davis, CA. Accessed January 19, 2011 https://www.geology.ucdavis.edu/~shlemonc/trips/SantaCruz_10/fieldguide.htm.

van der Meer J.W. 1988. Rock Slopes and Gravel Beaches Under Wave Attack, Ph.D.

dissertation, Delft University of Technology, The Netherlands. (Also Delft Hydraulics

Publication No.396).

APPENDIX A

Inundation Levels for Sea Level Rise Scenarios

Technical Report Appendix A, Page A-1

Figure A-1. Inundation limit of MHHW tide level, year 2010

Technical Report Appendix A, Page A-2

Figure A-2. Inundation limit of MHHW tide level, year 2050

Technical Report Appendix A, Page A-3

Figure A-3. Inundation limit of MHHW tide level, year 2100

Technical Report Appendix A, Page A-4

Figure A-4. Inundation limit of MHHW tide level with effects of 100-year wave storm, year 2010

Technical Report Appendix A, Page A-5

Figure A-5. Inundation limit of MHHW tide level with effects of 100-year wave storm, year 2050

Technical Report Appendix A, Page A-6

Figure A-6. Inundation limit of MHHW tide level with effects of 100-year wave storm, year 2100