ERDC/CHL TR-20-5 

Alabama Barrier Island Restoration 

Assessment Life-Cycle Structure Response 

Modeling

Coastal and Hydraulics Laboratory 

Victor M. Gonzalez, Fabian A. Garcia-Moreno, Jeffrey A. Melby, 

Norberto C. Nadal-Caraballo, and Elizabeth S. Godsey

April 2020

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 ERDC/CHL TR-20-5 

April 2020 

Alabama Barrier Island Restoration Assessment 

Life-Cycle Structure Response Modeling

Victor M. Gonzalez, Fabian A. Garcia-Moreno, Jeffrey A. Melby, 

and Norberto C. Nadal-Caraballo

Coastal and Hydraulics Laboratory

US Army Engineer Research and Development Center

3909 Halls Ferry Road

Vicksburg, MS 39180-6199

Elizabeth S. Godsey

Mobile District

US Army Corps of Engineers

109 St Joseph St.

Mobile, AL 36602

Final report

Approved for public release; distribution is unlimited

Prepared for Mobile District, US Army Corps of Engineers

Mobile, AL 36602

Under Project No. 452808
ERDC/CHL TR-20-5 ii

Abstract

Dauphin Island, a barrier island off the coast of Alabama, plays an 

important role in the protection of the state’s coastal natural resources. In 

2011, the State of Alabama constructed a rubble mound berm across a 

2 km breach in the western end of the barrier island to prevent oil from the 

Deep Water Horizon oil spill from migrating into the Mississippi Sound. 

The breach, referred to as the Katrina Cut, was caused by Hurricane 

Katrina in 2005. The US Army Corps of Engineers and the US Geological 

Survey (USGS) performed the Alabama Barrier Island Restoration 

Assessment study to assess the current and future function of Dauphin 

Island and evaluate potential restoration measures. A Monte Carlo lifecycle structure response assessment of the Katrina Cut rubble mound 

structure was performed as part of this study by the US Army Engineer 

Research and Development Center, Coastal and Hydraulics Laboratory. 

Damage, wave transmission, and reliability were computed within the 

context of the decadal barrier evolution analysis performed by the USGS 

for various storminess and relative sea level change scenarios. The 

presence of a beach in front of the structure plays an important role in its 

protection. The breaching potential for measures was evaluated. 

DISCLAIMER: The contents of this report are not to be used for advertising, publication, or promotional purposes. 

Citation of trade names does not constitute an official endorsement or approval of the use of such commercial products. 

All product names and trademarks cited are the property of their respective owners. The findings of this report are not to 

be construed as an official Department of the Army position unless so designated by other authorized documents.

DESTROY THIS REPORT WHEN NO LONGER NEEDED. DO NOT RETURN IT TO THE ORIGINATOR.
ERDC/CHL TR-20-5 iii

Contents

Abstract.................................................................................................................................... ii

Figures and Tables..................................................................................................................iv

Preface...................................................................................................................................viii

1 Introduction ......................................................................................................................1

1.1 Background........................................................................................................1

1.2 Objectives...........................................................................................................4

1.3 Approach ............................................................................................................4

2 Dauphin Island Katrina Cut Structure and Restoration Measures............................5

2.1 Katrina Cut structure.........................................................................................5

2.2 Restoration measures.......................................................................................8

3 Probabilistic Storm Forcing..........................................................................................14

3.1 Forcing data .....................................................................................................14

3.2 Monte Carlo storm sampling...........................................................................17

3.3 Storm sampling implementation for morphological model...........................19

4 Life-Cycle Structure Response Modeling....................................................................22

4.1 Armor stability and reliability ......................................................................... 22

4.2 Storm peak reliability analysis........................................................................24

4.3 Damage progression ...................................................................................... 28

4.3.1 Damage progression sensitivity analysis.................................................................28

4.3.2 Incorporation of uncertainty .....................................................................................34

4.3.3 Damage progression for restoration measures ......................................................35

4.3.4 Damage progression analysis without storminess bins .........................................40

5 Wave Transmission ........................................................................................................43

5.1 Wave transmission analysis........................................................................... 43

5.2 Impact of restoration measure 5 on transmitted wave energy ................... 45

6 Summary and Conclusions ...........................................................................................49

References.............................................................................................................................51

Appendix A: Stability and Damage Analysis......................................................................54

Appendix B: Wave transmission..........................................................................................60

Appendix C: Wave Transmission Analysis Results (Figures) ...........................................63

Acronyms and Abbreviations...............................................................................................79

Report Documentation Page
ERDC/CHL TR-20-5 iv

Figures and Tables

Figures

Figure 1-1. Aerial images of Katrina Cut area in Dauphin Island from 2000 to 

2019................................................................................................................................................3

Figure 2-1. Typical section design drawing for Katrina Cut structure. (Permit No. 

SAM-2010-00822-JBE).................................................................................................................6

Figure 2-2. Damage area in Katrina Cut rubble mound structure identified 

in 2016 survey................................................................................................................................7

Figure 2-3. Leeside damage at damage area 3........................................................................7

Figure 2-4. Alignment of pipelines that cross underneath the Katrina Cut 

structure..........................................................................................................................................8

Figure 2-5. R4: West and East End beach and dune nourishment........................................9

Figure 2-6. R5: Back Barrier tidal flats and marsh habitat restoration..................................9

Figure 2-7. R6: West End beach and dune nourishment.......................................................10

Figure 2-8. R7: West End and Katrina Cut beach and dune nourishment. .........................10

Figure 3-1. Map Showing portions of ADCIRC mesh covering Mobile and 

Pensacola Bay. .............................................................................................................................15

Figure 3-2. Map showing SWAN wave model domains. .........................................................16

Figure 3-3. FEMA Region IV save points and SWAN model region 2....................................17

Figure 3-4. General workflow for computing structure response and reliability 

for Katrina Cut structure.............................................................................................................18

Figure 3-5. Histogram showing the distribution of number of TCs across the 

realizations for each storminess bin. ........................................................................................21

Figure 4-1. DEM for determining the beach profile (10X vertical exaggeration).................25

Figure 4-2. Seaside damage variation surface for ST1..........................................................30

Figure 4-3. Leeside damage variation surface for ST1. .........................................................30

Figure 4-4. Seaside damage variation surface for ST2..........................................................32

Figure 4-5. Leeside damage variation surface for ST2. .........................................................32

Figure 4-6. Seaside damage variation surface for ST3..........................................................34

Figure 4-7. Leeside damage variation surface for ST3...........................................................34

Figure 4-8. Profile line for toe depth selection at known breaching point. ..........................36

Figure 4-9. Profiles extracted from ST3SL 3 DEMs for R0, R4, R5, R6, and R7 

restoration measures at the end of 10-year simulation.........................................................37

Figure 5-1. Schematic diagram of wave transmission analysis configurations..................44

Figure 5-2. Line for profile extraction of water depths and footprint of restoration 

measure 5 on the leeside of the Katrina Cut structure..........................................................46

Figure 5-3. Comparison between restoration measure 5 profiles and existing 

profile (NAVD88). .........................................................................................................................46
ERDC/CHL TR-20-5 v

Figure 5-4. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and SLR configurations with a crest height of 

2.44 m. Leeside depth: -0.47 m................................................................................................47

Figure 5-5. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and SLR configurations with a crest height of 

2.44 m. Leeside depth: -0.47 m................................................................................................47

Figure 5-6. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and SLR configurations with a crest height of 

2.44 m. Leeside depth: -0.47 m................................................................................................48

Figure A-1. Illustration of rubble mound damage parameters..............................................56

Figure B-1. Wave transmission coefficient for SB 2 using Equation B-2. ............................62

Figure B-2. Wave transmission coefficient for SB 2 using Equation B-6. ............................62

Figure C-1. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: -0.47 m...............................................................................................63

Figure C-2. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: -0.47 m...............................................................................................63

Figure C-3. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: -0.47 m...............................................................................................64

Figure C-4. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: -0.47 m...............................................................................................64

Figure C-5. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: -0.47 m...............................................................................................65

Figure C-6. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: -0.47 m...............................................................................................65

Figure C-7. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: -0.47 m...............................................................................................66

Figure C-8. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: -0.47 m...............................................................................................66

Figure C-9. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: -0.47 m...............................................................................................67

Figure C-10. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: -0.19 m...............................................................................................67

Figure C-11. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: -0.19 m...............................................................................................68
ERDC/CHL TR-20-5 vi

Figure C-12. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: -0.19 m...............................................................................................68

Figure C-13. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: -0.19 m...............................................................................................69

Figure C-14. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: -0.19 m...............................................................................................69

Figure C-15. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: -0.19 m...............................................................................................70

Figure C-16. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: -0.19 m...............................................................................................70

Figure C-17. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: -0.19 m...............................................................................................71

Figure C-18. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: -0.19 m...............................................................................................71

Figure C-19. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: 0.38 m................................................................................................72

Figure C-20. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: 0.38 m................................................................................................72

Figure C-21. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST1 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: 0.38 m................................................................................................73

Figure C-22. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: 0.38 m................................................................................................73

Figure C-23. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: 0.38 m................................................................................................74

Figure C-24. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST2 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: 0.38 m................................................................................................74

Figure C-25. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

0.91 m. Leeside depth: 0.38 m................................................................................................75

Figure C-26. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

1.83 m. Leeside depth: 0.38 m................................................................................................75
ERDC/CHL TR-20-5 vii

Figure C-27. Mean maximum transmitted waves and wave breaking classification 

at leeside for all ST3 toe depths and RSLC configurations with a crest height of 

2.44 m. Leeside depth: 0.38 m................................................................................................76

Figure C-28. Restoration measure 5 modeling conditions and leeside bottom 

influence on transmitted waves for 0.91 m crest height. ......................................................77

Figure C-29. Restoration measure 5 modeling conditions (after 10 years) and 

leeside bottom influence on transmitted waves for 1.81 m crest height. ...........................77

Figure C-30. Restoration measure 5 modeling conditions and leeside bottom 

influence on transmitted waves for 2.44 m crest height. ......................................................78

Tables

Table 2-1. Class 5 riprap gradation and corresponding nominal stone size.........................6

Table 3-1. Storminess bin classification and associated probability....................................20

Table 3-2. Comparison of storm count frequency and Poisson distribution 

probability for one thousand 10-year life-cycle realizations...................................................21

Table 4-1. Mean reliability values per storminess bin and RSLC conditions for 

rubble mound structure stability................................................................................................26

Table 4-2. Percent difference in reliability with respect to storminess bin 1.......................27

Table 4-3. Percent difference in reliability with respect to RSLC scenario 1. ......................27

Table 4-4. Standard deviation of the computed reliability values.........................................27

Table 4-5. Seaside and leeside damage for storminess bin 1..............................................29

Table 4-6. Seaside and leeside damage for storminess bin 2..............................................31

Table 4-7. Seaside and leeside damage for storminess bin 3. .............................................33

Table 4-8. Table of statistical quantities where N (mean, standard deviation) 

denotes normally distributed parameter and D (value) represents a deterministic 

value. .............................................................................................................................................35

Table 4-9. Damage progression for STSL3 restoration measures representative 

toe depths.....................................................................................................................................38

Table 4-10. Accumulated mean seaside damage and confidence limits for 

restoration measures at the end of 10-year simulation period.............................................38

Table 4-11. Accumulated mean leeside damage and confidence limits for 

restoration measures at the end of 10-year simulation period.............................................39

Table 4-12. Comparison of computed reliability using peaks and storm 

hydrograph....................................................................................................................................39

Table 4-13. Seaside damage comparison between no-storminess bin and 

storminess bin approaches........................................................................................................41

Table 5-1. Mean maximum transmitted waves for all storminess bins, toe depths, 

and RSLC configurations with a crest height of 2.44 m.........................................................44
ERDC/CHL TR-20-5 viii

Preface

This study was conducted for the US Army Corps of Engineers, Mobile 

District, under Project No. 452808. The Program Manager for the study

was Mr. Justin S. McDonald.

The work was performed by the Harbors, Entrances, and Structures

Branch of the Navigation Division, US Army Engineer Research and 

Development Center, Coastal and Hydraulics Laboratory (ERDC-CHL). At 

the time of publication of this report, Mr. James D. Gutshall was Chief, 

Harbors, Entrances, and Structures Branch; Dr. Jacqueline S. Pettway was 

Chief, Navigation Division; and Mr. Charles E. Wiggins was the Technical 

Director for Navigation. The Deputy Director of ERDC-CHL was 

Mr. Jeffrey R. Eckstein, and the Director was Dr. Ty V. Wamsley.

Many of the methods and tools applied in this study were developed as 

part of the Civil Works Monitoring Completed Navigation Projects 

(MCNP) Work Unit Reliability of Rubble Mound Structures, of which 

Dr. Lyndell Z. Hales is Program Manager.

The Commander of ERDC was COL Teresa A. Schlosser, and the Director 

was Dr. David W. Pittman.
ERDC/CHL TR-20-5 1

1 Introduction

Dauphin Island, a barrier island off the coast of Alabama, plays an 

important role in the protection of the state’s coastal natural resources. 

During an 8-year span from 2004 to 2012, the island experienced 

significant tropical storm events and a man-made disaster that affected 

the morphology and ecology of the island. These included Hurricanes 

Ivan, Katrina, and Isaac, as well as the Deep Water Horizon oil spill. One 

of the most significant impacts was a 2 km1,2 breach in the barrier island 

created by Hurricane Katrina in 2005. As a response to the 2010 Deep 

Water Horizon oil spill, a temporary rubble mound structure was 

constructed across the cut to prevent oil moving into the Mississippi 

Sound. This technical report presents the results of the Katrina Cut rubble 

mound structure response modeling that was part of the wider Alabama 

Barrier Island Restoration Assessment Study sponsored by the National 

Fish and Wildlife Foundation and the State of Alabama. The structure 

response modeling was performed by the Coastal and Hydraulics 

Laboratory, US Army Engineer Research and Development Center. 

Structure damage and reliability estimates were computed for different 

storminess and sea level change scenarios for select restoration measures. 

Wave transmission due to overtopping was also evaluated.

1.1 Background

The passage of Hurricane Katrina west of Dauphin Island, AL produced a 

2 km wide breach in the unpopulated western end of the barrier island. 

This breach is referred to as the Katrina Cut. Breaching and natural closing 

of the barrier island have been documented in the past, with closures of 

large breaches occurring over decadal time scales (USGS et al. 2017). 

Nevertheless, to protect the Mississippi Sound following the 2010 Deep 

Water Horizon oil spill, the Katrina Cut was closed with a rubble mound 

structure. Figure 1-1 shows the changes at the Katrina Cut area from the 

 

1 For a full list of the spelled-out forms of the units of measure used in this document, please refer to US 

Government Publishing Office Style Manual, 31st ed. (Washington, DC: US Government Publishing 

Office 2016), 248-52, https://www.govinfo.gov/content/pkg/GPO-STYLEMANUAL-2016/pdf/GPOSTYLEMANUAL-2016.pdf.

2 For a full list of the unit conversions used in this document, please refer to US Government Publishing 

Office Style Manual, 31st ed. (Washington, DC: US Government Publishing Office 2016), 345-7, 

https://www.govinfo.gov/content/pkg/GPO-STYLEMANUAL-2016/pdf/GPO-STYLEMANUAL-2016.pdf.
ERDC/CHL TR-20-5 2

year 2000 to the year 2019, including structure construction. The 2005 

image predates Hurricane Katrina and shows a smaller breach caused by 

Hurricane Ivan the year prior. In the years following the closure of the 

Katrina Cut, a beach formed on the Gulf of Mexico (south) side of the 

structure. The State of Alabama engaged the US Geological Survey (USGS)

and the US Army Corps of Engineers (USACE) to perform a study, through 

a grant from the National Fish and Wildlife Foundation, Gulf 

Environmental Benefit Fund, to investigate feasible alternatives for the 

sustainable restoration of Dauphin Island for habitat and species impacted 

by the 2010 Deepwater Horizon oil spill. 
ERDC/CHL TR-20-5 3

Figure 1-1. Aerial images of Katrina Cut area in Dauphin Island from 2000 to 2019. 

 

 

 

 
ERDC/CHL TR-20-5 4

1.2 Objectives

The Alabama Barrier Island Restoration Assessment project’s main 

objective was to assess the current and future function of Dauphin Island, 

develop options for restoration, and evaluate the feasibility and cost 

associated with restoration actions. The focus of the restoration actions 

was the optimization of the island's resiliency to storm events, the 

enhancement of wildlife habitat, and increase of sustainability. The 

Katrina Cut rubble mound structure was intended to be a temporary 

measure to prevent oil migration into the Mississippi Sound from the 

Deep Water Horizon oil spill. The structure is still present and therefore 

evaluated as part of the study. 

The structural response modeling of the Katrina Cut structure was 

performed to determine damage, wave transmission, and reliability within 

the context of the decadal modeling performed for the restoration 

measures. Specific objectives of this study included the following: 

• Characterization of the wave and water level forcing for the 

morphologic and the rubble mound structure response modeling for 

selected restoration measures. 

• Evaluation of the performance of the Katrina Cut rubble mound 

structure for restoration measures for varying storminess conditions 

and relative sea level change (RSLC) scenarios.

1.3 Approach

A Monte Carlo life-cycle analysis of the rubble mound structure response 

was applied using a 10-year simulation period consistent with the decadal 

modeling approach used by the USGS in its barrier island evolution 

forecasting (Mickey et al. 2020). Synthetic storms developed for a regional 

joint probability method coastal hazard storm surge study were sampled to 

generate the forcing for the 10-year simulation period. Rubble mound 

response was evaluated for the range of probable beach response 

scenarios. The outcome of the rubble mound structure assessment will 

inform the costing of restoration measures and aid decision makers and 

stakeholders in determining the fate of the temporary rubble mound 

structure at the Katrina Cut. 
ERDC/CHL TR-20-5 5

2 Dauphin Island Katrina Cut Structure and 

Restoration Measures

A small breach, formed through the western portion of Dauphin Island, 

AL, after hurricane Ivan, was significantly widened to a 2 km wide breach 

by Hurricane Katrina the following year. This breach was closed in 2011 

with a rubble mound berm by the State of Alabama in response to the 

2010 Deep Water Horizon oil spill. The original objective was to build a 

temporary berm that would prevent oil migration from the spill into the 

Mississippi Sound (USGS et al. 2017). 

2.1 Katrina Cut structure

The Katrina Cut structure was designed as a temporary structure. The 

permit for the structure includes drawings and specifications as well as the 

performance criteria on which the design was based. The typical cross

section is shown in Figure 2-1. It consists of a 2:1 side-slope rubble mound 

structure with a sand core surrounded by a geotextile filter fabric. The top 

of the core was designed to an elevation of 1.52 m North American Vertical 

Datum of 1988 (NAVD88), the estimated wave run-up elevation under 

normal conditions. The definition of normal was not found, but the use of 

the term seems consistent with the intent of it being a temporary 

structure. The design crest height was 2.44 m NAVD88, with a 0.91 m 

thick armor top layer. The armor design specified the use of Class 5 riprap 

as defined in specification 814 of the 2008 Alabama Standard 

Specifications for Highway Construction. Class 5 stone is composed of 

nearly rectangular quarry stone with the size distribution shown in 

Table 2-1. The median size for this stone was 450 kg, which, according to 

the design, was estimated to provide sufficient protection against 1.37 m 

wave heights. The design section also included a 12 in. #57 choke stone 

layer on top. A choke stone is a small-sized stone used to stabilize the 

surface of an open-graded material. Based on field assessments of the 

structure conducted in September 2016, it is unclear if the choke stone was 

placed. It may have been placed and washed away after installation. This 

study assumes that the choke stone does not have any effect on the 

performance of the structure. The structure had a crest width of 6.10 m, 

with a toe elevation of -2.44 m at the deepest part of the breach. 
ERDC/CHL TR-20-5 6

Figure 2-1. Typical section design drawing for Katrina Cut structure. 

(Permit No. SAM-2010-00822-JBE) 

Table 2-1. Class 5 riprap gradation and corresponding nominal 

stone size. 

Percent passing Weight 

(kg) 

Dn 

(m) 

90% 900 0.7 

50% 450 0.55 

25% 100 0.33 

The nominal median stone size (Dn50), also referred to as the equivalent 

cube length of the median stone, was computed assuming a specific armor 

weight, γγss, of 2,643 kg/m3 using Equation 2-1. The computed median 

nominal stone size was 0.55 m. 

DDnn50 = �

WW50

γγss

�

1

3�

(2-1)

An assessment of the aerial images shows that over time the structure 

has suffered damage at some locations since its construction. This was 

confirmed by a 2016 survey and inspection commissioned by the USACE 

Mobile District. The survey revealed multiple areas where the crest 

elevation had been lowered from the original design elevation. Figure 2-2 

shows an aerial image with six damage areas identified in 2016. 

Figure 2-3 presents a photograph of leeside damage at damage area 3. 

Gas and/or oil pipelines run underneath the west end of the Katrina Cut. 

Two of the alignments run underneath damage areas 1 and 2. The design 

of the structure contemplated the presence of pipelines and called for the 

installation of a structural geogrid under the berm at locations were they 

cross. Figure 2-4 shows a schematic of the location of the pipeline 

alignments, represented as yellow lines running perpendicular to the 

rubble mound structure. 
ERDC/CHL TR-20-5 7

Figure 2-2. Damage area in Katrina Cut rubble mound structure identified 

in 2016 survey. 

Figure 2-3. Leeside damage at damage area 3. 
ERDC/CHL TR-20-5 8

Figure 2-4. Alignment of pipelines that cross 

underneath the Katrina Cut structure. 

2.2 Restoration measures

The life-cycle structure response assessment was performed in the context 

of restoration measures proposed for Dauphin Island. Seven restoration 

measures and a no-action measure were implemented by the USGS for 

initialization of the barrier island response modeling over a 10-year 

timeframe. The description of the restoration measures is presented in 

Table 2-2. The information in this section summarizes the analysis of the 

restoration measures that were documented in Mickey et al. (2020).

 Table 2-2. Description of restoration measures 

Restoration 

Measure Restoration Description 

R0 No Action

R1 Katrina Cut – sand berm

R2 Pelican Island Southeast nourishment

R3 Sand Island platform nourishment and sand bypassing

R4 West and East End beach and dune nourishment

R5 Back Barrier tidal flats and marsh habitat restoration

R6 West End beach and dune nourishment

R7 West End and Katrina Cut beach and dune nourishment
ERDC/CHL TR-20-5 9

Not all restoration measures impact the area of Katrina Cut. Measure R1 

was a sensitivity run where the Katrina rubble mound structure was 

replaced with a sand berm. Measures R2 and R3 are implemented to the 

southeast of the Katrina Cut at Pelican Island and the Sand Island 

platform and do not produce an impact at the study area. Restoration 

measures R4, R5, R6, and R7 are implemented in the vicinity of the 

Katrina Cut area. Measure R4 includes nourishment of the beach east of 

the Katrina Cut and the shorefront area east of Pelican Island. Figure 2-5 

shows the shoreline comparison between R0 (blue line) and R4 (orange 

line). Measure R5 refills borrow pits at the back of the Katrina Cut area 

and restores the back-bay marsh and island platforms there. The 

comparison of the no-action shoreline and this measure is shown in 

Figure 2-6. Measure R6 represents a modification of R4 that moves the 

restored dune area on the western side of Katrina Cut landward to better 

align with locations observed in 1950s photographs (Mickey et al. 2020). It 

does not include the beach nourishment in the east. Figure 2-7 shows a 

comparison of the shorelines for Measure R7 and R0. Measure R7 extends 

the beach restoration in R6 to the front of the Katrina Cut structure, and 

the shoreline is shown in Figure 2-8. 

Figure 2-5. R4: West and East End beach and dune nourishment. 

Figure 2-6. R5: Back Barrier tidal flats and marsh habitat restoration. 
ERDC/CHL TR-20-5 10

Figure 2-7. R6: West End beach and dune nourishment. 

Figure 2-8. R7: West End and Katrina Cut beach and dune nourishment. 

The modeling of the barrier island restoration measures was done for 

three levels of tropical cyclone (TC) frequency and intensity, which will be 

referred to herein as storminess levels. A Monte Carlo sampling method 

was used to sample 1,000 life cycles of TCs. This method is described in 

Chapter 3. The USGS (Mickey et al. 2020) developed a method to group 

the life-cycle storm sequences or realizations into storminess bins (ST)

based on the corresponding island response using a one-dimensional (1D)

proxy model framework (CA1D). Additional description of the ST

approach can be found in section 3.3. A single sequence of storms was 

selected per ST to be modeled in the full modeling framework. The 

framework consists of Delft3D (Deltares 2019) and the empirical dune 

growth model (EDGR), along with numerical model XBeach (Roelvink et 

al. 2009), to simulate storm-induced erosion. Three 10-year ST were run, 

each adopting three RSLC scenarios. These were determined using a 

relative rate of sea level change (SLC) of 3.61 mm/year along with the 

USACE SLC curve calculator (version 2017.55) for low, intermediate, and 

high curves at Dauphin Island National Oceanic and Atmospheric 

Administration gage 8735180 (Mickey et al. 2020). These relative sea level 

change cases (SL), projected to 2070, are as follows:

• SL1: 0.3 m increase to water level at offshore boundary 

• SL2: 0.5 m increase to water level at offshore boundary

• SL3: 1.0 m increase to water level at offshore boundary.
ERDC/CHL TR-20-5 11

The full model framework was initialized with the 2015 Dauphin Island 

configuration. It was run for two scenarios based on the combination ST 

and SL discussed above. Scenarios of storminess bin 2 with SL1 (ST2SL1) 

and storminess bin 3 with SL3 (ST3SL3) were run for the restoration 

measures R2 through R7. R1 was run only with ST3SL3 and resulted in a 

breach in the Katrina Cut area 3.5 km wide (Mickey et al. 2020). Figures 

2-9 to 2-18 show the comparison of the initial and resulting shorelines of 

the USGS (Mickey et al. 2020) analysis. The figures provide a comparison 

of changes in Dauphin Island shoreline from initiation at T=0 year to the 

end of the modeled period at T=1o years for the restoration measures and 

two scenarios run. Each figure contains an inset with a zoom of the Katrina 

Cut area. These results were considered in the damage progression 

modeling of the rubble mound structure. 

Figure 2-9. R0: No action for T0 to T10 for ST2SL1. 

Figure 2-10. R0: No action for T0 to T10 for ST3SL3. 

Figure 2-11. R4: West and East End beach and dune nourishment T0 to T10 

for ST2SL1. 
ERDC/CHL TR-20-5 12

Figure 2-12. R4: West and East End beach and dune nourishment T0 to T10 

for ST3SL3. 

Figure 2-13. R5: Back Barrier tidal flats and marsh habitat restoration for T0 to T10 

for ST2SL1. 

Figure 2-14. R5: Back Barrier tidal flats and marsh habitat restoration for T0 to T10 

for ST3SL3. 

Figure 2-15. R6: West End beach and dune nourishment for T0 to T10 for ST2SL1. 
ERDC/CHL TR-20-5 13

Figure 2-16. R6: West End beach and dune nourishment for T0 to T10 for ST3SL3. 

Figure 2-17. R7: West End and Katrina Cut beach and dune nourishment T0 to T10 

for ST2SL1. 

Figure 2-18. R7: West End and Katrina Cut beach and dune nourishment T0 to T10 

for ST3SL3. 
ERDC/CHL TR-20-5 14

3 Probabilistic Storm Forcing

Tropical cyclones represent the main driver of coastal hazards in the Gulf 

of Mexico coastline of the United States. As previously described, the 

Katrina Cut in Dauphin Island, AL, was caused by Hurricanes Ivan and 

Katrina in back-to-back years. The forcing for the hydrodynamic and 

morphological modeling, as well as the structural response modeling for 

Dauphin Island consisted of sampled TCs using a Monte Carlo Life-Cycle 

approach. 

3.1 Forcing data

The statistical assessment of coastal storm hazards in hurricane-prone 

coastal areas requires the development of a joint probability model of TC 

forcing parameters. The joint probability method with optimal sampling 

(JPM-OS) became the standard joint probability analysis model for the 

assessment of coastal storm hazards for hurricane-prone US coastlines 

following Hurricane Katrina (Resio et al. 2007; Nadal-Caraballo et al. 

2015). In this method, synthetic tropical storms that cover forcing 

probability space are developed from a joint probability model of tropical 

storm parameters (i.e., reference location [xo], heading [θ], central 

pressure deficit [Δp], radius of maximum winds [Rmax], and translational 

speed [Vt] and modeled in high-resolution numerical physics-based 

models to obtain the corresponding response. The annual exceedance 

probability of coastal storm hazards at a given site is a function of three 

main components: the storm recurrence rate, the joint probability of 

characteristic TC forcing parameters, and the TC responses (e.g., storm 

surge, wave height). 

The forcing used for this study was based on comprehensive coastal storm 

modeling that was completed for coastal Alabama under Federal 

Emergency Management Agency (FEMA) Region IV Risk Mapping, 

Analysis and Planning study and Digital Flood Insurance Rate Map update 

for the Florida Panhandle and Alabama coasts. The modeling included 

waves and water levels for 295 synthetic tropical storms making landfall in 

the region that were derived using JPM-OS (URS/Dewberry 2012).

Water levels and waves were computed using two different models: (1) the 

Simulating WAves Nearshore (SWAN) model (http://www.swan.tudelft.nl), used 

for producing offshore, regional, and coastal wave conditions, and (2) the 
ERDC/CHL TR-20-5 15

Advanced Circulation (ADCIRC) model (ADCIRC 2017; Luettich et al. 

1992; Westerink et al. 1992), which was used to simulate two-dimensional

depth-averaged surge and circulation responses to the storm conditions. 

The modeling used two different ADCIRC meshes, one that had coarser 

resolution covering much of the western North Atlantic Ocean, the 

Caribbean Sea, and the Gulf of Mexico and a finer resolution covering the 

floodplain areas within Mobile and Baldwin Counties, Alabama; and 

Escambia, Santa Rosa, Okaloosa, Walton, Bay and Gulf Counties, Florida 

(Figure 3-1). The SWAN model was run with three levels of grids, a basin, 

a nested regional, and then 5 coastal grids (Figure 3-2).

Figure 3-1. Map Showing portions of ADCIRC mesh covering Mobile and 

Pensacola Bay.1 

 

1 Massey C., T, E. Godsey, R. Jensen, M. Bryant, Y. Ding, and M. Owens. 2020. Unpublished draft. An 

Overview of the Coastal Storm Model Simulations of Waves and Water Levels for the Alabama Coastal 

Comprehensive Plan.
ERDC/CHL TR-20-5 16

Figure 3-2. Map showing SWAN wave model domains.1 

Save point locations were set up in the region for obtaining the ADCIRC 

results, and SWAN outputs were obtained at the grid nodes. Figure 3-3 

shows the save point locations for the ADCIRC modeling and the extents 

of the coastal region 2 SWAN grid that covers the project area. The red 

points denote landfall locations of the synthetic storms. Save point 291 was 

selected to extract water levels for the structure response modeling. Wave 

data was extracted from the closest SWAN grid node to save point 291. 

 

1 Massey C., T, E. Godsey, R. Jensen, M. Bryant, Y. Ding, and M. Owens. 2020. Unpublished draft. An 

Overview of the Coastal Storm Model Simulations of Waves and Water Levels for the Alabama Coastal 

Comprehensive Plan.
ERDC/CHL TR-20-5 17

Figure 3-3. FEMA Region IV save points and SWAN model region 2. 

3.2 Monte Carlo storm sampling

The forcing for the modeling of Dauphin Island morphological evolution 

and structure response consisted of a suite of TCs sampled for a 10-year 

simulation period and multiple life cycles. The number of synthetic TCs in 

a given year was computed using the Poisson distribution and the storm 

recurrence rate (SRR). The SRR corresponding to the location of save 

point 291 was computed as 9.63 × 10-4 storms/year/km using the Gaussian 

kernel methodology developed by Chouinard and Liu (1997) and 

Chouinard et al. (1997). Consistent with experiments performed by NadalCaraballo et al. (2019), an optimal kernel size of 200 km was applied to 

compute the SRR. The SRR was converted to storms per year by 

multiplying by 400 km to represent the recurrence rate associated with 

storms passing within 200 km of the coastal reference location (CRL) 

closest to save point 291. This results in a rate of 0.385 storms/year. 

The occurrence of TCs within the North Atlantic basin is often described as 

a Poisson process (e.g., Chouinard and Liu 1997; Elsner et al. 2001; 

Sommers 2003). The Poisson distribution can be used to compute the 

number of times an event occurs in an interval of time knowing the 

recurrence rate λ of X, where X is a random variable representing the 

number of TCs in a year. When applied to TCs, the probability of observing 

x TCs in an interval of time is given by

  !

x λ λ e PX x

x



  (3-1)
ERDC/CHL TR-20-5 18

The Poisson distribution was used to determine the number of storms in a 

given year within the Monte Carlo sampling (e.g., two storms in the first 

year of the simulation, zero storms the second, one storm the third). Storm 

sequences that sampled a total number of storms that matched the SRR 

were retained for the analysis. The variability considered in the sampled 

life cycles included the selection of different combinations of storms and 

parameter uncertainty. Synthetic TCs with landfall locations within 

200 km of the CRL were sampled randomly using their probability masses 

as sampling weights. 

The workflow for the Katrina structure response analysis is shown in 

Figure 3-4. The TC storms are sampled at save point 291, and 

corresponding wave conditions transformed to the toe of the structure 

where structure response is computed. The reliability of the structure is

then computed using the outcome of the simulations with respect to a 

limit state. Reliability criteria is described further in section 4.1. 

Figure 3-4. General workflow for computing structure response and reliability 

for Katrina Cut structure. 

291
ERDC/CHL TR-20-5 19

Since the Katrina Cut structure was built in relatively shallow water and 

the deposition of sand seaward of the structure has further decreased the 

water depth there, wave conditions are considered depth limited. A 

limitation to the zero-moment wave heights at the structure was applied 

based on Battjes and Janssen (1978), applying Equation 3-2 to obtain 

depth limited waves based on both depth and wave steepness:

HHmm0mmmmmm = 0.1LLLLLLLLh(kkkk) (3-2)

where HHmm0mmmmmm is the maximum depth limited significant wave height; L is 

the wavelength; k is the wave number; and d the water depth. This wavebreaking criteria is also applied in spectral-wave model STWAVE (Smith et 

al. 1999). Both k and L are based on the peak wave period and computed 

using the linear wave theory dispersion relation.

3.3 Storm sampling implementation for morphological model

Storm sampling for the CA1D modeling framework for the morphologic 

analysis was performed by the USGS (Mickey et al. 2020) using the

methodology described in section 3.2. It resulted in one thousand 20-year 

realizations of storm sequences. The first 10 years out of the 20-year 

periods were selected for the 1D analysis based on a change in the 

timeframe over which the measures needed to be evaluated. This 10-year 

simulation period is relevant for coastal planning and management and 

aligns with the overall objective of the Alabama Barrier Island Restoration 

Assessment. Each of the one thousand 10-year realizations was classified 

into three storminess bins (low, medium, and high) based on the island 

response obtained with the 1D model. Specifically, classification was based 

on the number of events that were estimated to result in overwash of the 

representative dune profile (Mickey et al. 2020). 

The number of life cycles per storminess bin is shown in Table 3-1. The 

Katrina Cut structure response modeling was performed for the same 

three storminess bins to harmonize the structure analysis with the 

Dauphin Island evolution modeling approach. However, as mentioned in 

section 2.2, only a single representative storm sequence per storminess bin 

was selected to perform the morphologic modeling associated with the full 

modeling framework for all RSLC scenarios. The structure response 

analysis was performed for all the life cycles since there was no 

computational restriction for running the simulations. However, the 

morphologic response used in the structure modeling corresponds to a 
ERDC/CHL TR-20-5 20

single instance per storminess bin and RSLC case. It is recognized that 

representing the morphological response by a single sequence of storms or 

iteration precludes the quantification of the aleatory uncertainty 

associated with that response.

Table 3-1. Storminess bin classification and associated probability. 

Storminess Bin Number of Life Cycles Probability 

1 157 13.3%

2 576 57.3%

3 267 29.4%

Initially, 20-year life cycles were computed. As discussed in section 3.2, 

the 1,000 life cycles were selected from a larger group based on how well 

each life cycle matched the target SRR. A tolerance of 0.05 storms/year 

was used as selection criteria. Sampled sequences with an SRR within this 

tolerance were selected. However, this SRR matching was not strictly 

adhered to in the final life cycles used because only the first 10 years were 

selected for forcing the decadal model. Figure 3-5 shows the histogram of 

the number of storms per 10-year period for each storminess bin, 

compared to the expected distribution of storms per 10-year period based 

on the Poisson distribution and an SRR equal to 0.385 storms/year 

(λ=3.85 storms/decade). The histogram is shown for the combined one 

thousand 10-year realizations. The frequency of storms in the sampled 

10-year one thousand life cycles is compared to the frequency of storms 

predicted by the Poisson distribution in Table 3-2. The resulting number 

of storms per life cycle overpredicts the Poisson distribution for 3, 4, and 5 

storms per decade and underpredicts for the rest. The probability 

associated with the occurrence of each storminess bin was computed for 

the 10-year realizations using the Poisson distribution probabilities and 

storm relative probabilities. This process is described further in Mickey et 

al. (2020). The resulting probability for each storminess bin was 13.3%, 

57.3%, and 29.4% for ST1, ST2, and ST3, respectively. These values were 

also listed in Table 3-1. 
ERDC/CHL TR-20-5 21

Figure 3-5. Histogram showing the distribution of number of TCs across the 

realizations for each storminess bin. 

Storminess Bin 1 Storms for 10-year 

period.

Storminess Bin 2 Storms for 10-year 

period.

Storminess Bin 3 Storms for 10-year 

period.

Storminess Bins 1 to 3 Storms for 10-year 

period.

Table 3-2. Comparison of storm count frequency and Poisson distribution 

probability for one thousand 10-year life-cycle realizations. 

Number of Storms 

per Life-Cycle Bin 

Number of 

Life-Cycles Count 

Frequency 

(Number of Life 

Cycles per 

Bin/1000) 

Poisson 

Probability 

(λ=3.85) 

0 5 0.005 0.021

1 40 0.04 0.082

2 144 0.144 0.157

3 257 0.257 0.202

4 279 0.279 0.195

5 178 0.178 0.150

6 75 0.075 0.096

7 19 0.019 0.053

8 5 0.003 0.026
ERDC/CHL TR-20-5 22

4 Life-Cycle Structure Response Modeling

The armor stability of the Katrina Cut rubble mound structure was 

analyzed based on storm peak armor damage and seaside and leeside 

armor damage progression. Four assessments were performed to 

characterize the Katrina Cut rubble mound structure response within the 

context of the Alabama Barrier Island Restoration Assessment Project. 

• Reliability analysis of stone armor stability based on storm peaks to 

evaluate the effect of a beach versus no beach in front of the structure. 

• Stone armor damage progression sensitivity analysis for various 

forcing, RSLC, and crest height conditions to define the damage 

performance surface. 

• Stone armor damage progression analysis using the forcing conditions 

and RSLC scenarios that coincide with the full barrier island evolution 

modeling of the restoration measures as described in section 2.2. 

• A damage progression analysis that samples TCs outside the context of 

storminess bins by applying 10,000 life cycles from the full set of TCs 

to corroborate the results of the binning approach.

4.1 Armor stability and reliability

The performance of the rubble mound structure was assessed using a 

reliability-based probabilistic approach. The equations used for the 

computation of seaside and leeside stone armor stability and damage 

progression used for these analyses are described in Appendix A. The 

general objective of the approach is to determine, for a particular 

performance mode, the probability of exceeding a specific level of response 

for the structure. The reliability approach applied herein was based on 

sampling the storm forcing and computing the probabilistic response 

explicitly using Monte Carlo simulations. Because there are a large number 

of forcing and response variables that have significant statistical variability 

and correlations, this method is usually significantly more accurate for this 

type of engineering analysis than other simplified approaches.

Reliability is defined in terms of a limit state that characterizes the failure 

surface of the system. The limit state can be formulated in terms of a 

performance function as follows:
ERDC/CHL TR-20-5 23

gg = RR − SS �

< 0 ff ffffffffff

= 0 ll ss (ff ffffffffff)

> 0 nn ff ffffffffff

(4-1)

where g is the performance of the system, R is the strength of the system, S

is the load on the system, and failure refers to the inability to meet the 

limit state criterion. The probability of an unsatisfactory response (e.g.,

failure) is given by 

PPff = PP(gg ≤ 0) (4-2)

and the reliability is given by

RRff=1-PPff (4-3)

Armor seaside and leeside stability and damage equations were 

rearranged into resistance (strength) and forcing (load) terms to conform 

to Equation 4-1. The limit state equations for seaside and leeside armor 

stability for storm peak forcing are described in Appendix A. The 

resistance terms of the equation includes the acceptable eroded area of 

the structure, stone size, and density. The forcing terms include wave 

parameters and steepness, number of waves, and water depth. 

The reliability or probability of satisfactory performance was computed as 

the percent of realizations that the structure does not exceed the limit 

state. This limit state was defined with respect to specific levels of damage, 

S, at the end of a life cycle. In the following sections, damage is described 

by S = Ae/(Dn50)2 = normalized eroded area, Ae = eroded area and Dn50 the 

nominal stone size. The limit state is meant to describe acceptable versus 

non-acceptable performance. Herein, unacceptable performance is defined 

as breaching given the temporary nature of the structure that is being 

examined. Damage is computed for every iteration and compared against 

the limit state. The reliability is defined by the ratio of iterations for which 

the limit state is not exceeded and the total number of iterations. 

Melby and Kobayashi (2011) suggest that damage S = 4.5Nod, where Nod is 

the number of stones displaced, and this provides a practical description of 

damage. Damage levels of S = 0 – 5 are considered no damage, 5 to 20 are 

serviceable, and over 20 are near the ultimate limit state where holes in 

the armor layer would be exposed to the core and the structure would be 
ERDC/CHL TR-20-5 24

near the initiation of breaching. This is based on Melby (1999) who 

suggested that this damage level corresponded to exposure of the core 

material. However, this is somewhat heuristic and uncertain, particularly 

for this low-crested structure that is semi-protected by a beach. Physical 

modeling would yield a more accurate estimate of the failure limit state for 

damage for this structure, but that was not feasible within the scope of this 

project. For seaside stability, a maximum damage value of S=30 was 

considered to produce a breach in the structure. Leeside damage for a lowcrested structure such as this one can progress more rapidly than seaside. 

Plunging overtopping wave action on the mostly unrestrained leeside 

armor stones just below the leeside corner of the crest can create a notch 

that concentrates damage and produces rapid erosion that can progress 

towards the seaside of the structure forming a breach (Melby et al. 2015). 

For leeside damage, a more conservative damage level of S=15 was 

considered as breaching. 

4.2 Storm peak reliability analysis

An initial reliability analysis for structure stability was performed based on 

storm peaks for all combinations of storminess bins and RSLC conditions. 

Two beach configurations were considered that bracketed the variation of 

morphologic changes. These were a “with-beach” and “no-beach” condition. 

The no-beach condition was modeled using a toe elevation of -2.44 m. This 

represents the deepest point in the rubble mound structure alignment 

according to the design documents included in the permit. To determine a 

toe elevation representative of the beach at the initiation of the modeling 

(2015 conditions), a higher-resolution digital elevation model (DEM) was 

created using some of the same topobathy data sources used for the barrier 

island evolution modeling DEM. These data sources included the USGS post 

Isacc 2012 Lidar, USGS 2015 Barrier Island Lidar, USACE/USGS August 

2015, and the USACE September 2016 topo/bathy datasets. A threedimensional rendering of the higher-resolution DEM is shown in Figure 4-1, 

along with the location of transects used to evaluate different beach widths 

in front of the structure. Transect XS 3 represented the 2015 location with 

the smallest beach width seaward of the structure. A toe elevation of 0.49 m 

was extracted to represent the with-beach condition. 
ERDC/CHL TR-20-5 25

Figure 4-1. DEM for determining the beach profile (10X vertical exaggeration). 

For the reliability analysis of seaside and leeside stability, exceedance or 

non-exceedance of the limit state was computed for each storm in a life 

cycle, reliability computed per life cycle, and results grouped by storminess 

bin. The number of life cycles per bin was as listed in Table 3-2. A mean 

reliability value was computed for each storminess bin for each RSLC

scenario as shown in Table 4-1. The outcome of the reliability analysis 

quantified heuristic expectations of the structure response for varied beach 

widths and RSLC scenarios. Reliability is high for a wide beach and low for 

a narrow beach, which describes the fact that the structure is not 

vulnerable to breaches with a wide beach but will breach if no fronting 

beach is present. The effect of the beach is to increase the bottom elevation 

at the toe of the structure and limit the maximum wave height that 

impacts the structure. Reliability values generally decrease with an 

increase in storminess conditions. Also, as anticipated, reliability values 

decreased with an increase in sea levels. 

The presence of the beach reduces the wave energy that reaches the 

structure and results in higher reliability values. For the no-beach 

condition, the reliability of the structure varies from 0.75 for ST1SL1 to 

0.33 for ST3SL3. For the with-beach condition, the reliability varies from 1 

to 0.79. The largest decrease in reliability due to storminess bin occurs 

between ST1 to ST2. There is a small decrease in reliability from 

storminess bin ST2 to ST3 when compared with ST1. The results also show 

the degree to which the reliability of the structure is sensitive to rising 
ERDC/CHL TR-20-5 26

relative sea levels. The percent differences across ST and RSLC scenarios

were computed and are shown in Tables 4-2 and 4-3, respectively. As seen 

in Table 4-2, for the no-beach condition, the reduction in reliability from 

ST1 to ST2 for all RSLC conditions is in the order of 30%. The difference is 

between 4.1% and 11% for the with-beach condition. The reduction in 

reliability due to an increase in the RSLC for the no-beach condition is 

from 6.1% to 10.8% from SL1 to SL2 and 28.6% to 31.3% from SL1 to SL3

as seen in Table 4-3. For the with-beach condition, the maximum 

reduction in reliability from SL1 to SL2 is approximately 1% to 9.7% from 

SL1 to SL3. 

This initial analysis did not apply empirical parameter or forcing 

uncertainty in order to isolate the effects in the variation in storminess bin 

and RSLC. Nevertheless, the analysis does consider the sampling 

variability associated with the storm selection for the different life cycles. 

Since, as seen in Figure 3-3, a relatively low number of storms is sampled 

in each life cycle (storms are relatively infrequent) and the number of life 

cycles varies from 167 for ST1 to 576 for ST2, there is significant variability 

in the computed reliability for each life cycle. As mentioned earlier, the 

above reliability values were mean values for each storminess bin. 

Table 4-4 shows the standard deviation of the computed reliability values 

per each storminess bin and RSLC scenario. The spread of the results is 

larger for the no-beach condition than for the with-beach condition. 

Table 4-1. Mean reliability values per storminess 

bin and RSLC conditions for rubble mound 

structure stability. 

Reliability - No Beach

Case SL1 SL2 SL3 

ST1 0.75 0.67 0.53

ST2 0.49 0.46 0.35

ST3 0.48 0.44 0.33

Reliability – With Beach

ST1 1.00 1.00 0.97

ST2 0.94 0.90 0.84

ST3 0.93 0.90 0.79
ERDC/CHL TR-20-5 27

Table 4-2. Percent difference in reliability with respect 

to storminess bin 1. 

No Beach

 SL1 SL2 SL3

ST1 - - -

ST2 -34.7 -31.3 -34.0

ST3 -36.0 -34.3 -37.7

Beach

ST1 - - -

ST2 -4.1 -5.0 -11.0

ST3 -5.2 -6.0 -14.4

Table 4-3. Percent difference in reliability with respect 

to RSLC scenario 1. 

No Beach

 SL1 SL2 SL3 

ST1 - -10.8 -29.3

ST2 - -6.1 -28.6

ST3 - -8.3 -31.3

Beach

ST1 - 0 0

ST2 - -1.0 -7.2

ST3 - -0.8 -9.7

Table 4-4. Standard deviation of the computed 

reliability values. 

Reliability - No Beach

Case SL1 SL2 SL3 

ST1 0.3 0.33 0.36

ST2 0.3 0.28 0.26

ST3 0.24 0.23 0.2

Reliability – With Beach

ST1 0 0 0.02

ST2 0.12 0.14 0.19

ST3 0.10 0.10 0.15
ERDC/CHL TR-20-5 28

Collectively, the reliability values suggest that the structure would be judged 

as failed (breached) under the no-beach condition for all storminess bins 

and all RSLC scenarios. However, under the with-beach condition, the 

structure would be intact under all scenarios with no damage ST1 and would 

have serviceable damage under ST2 and ST3 scenarios.

4.3 Damage progression

Structure response over time was assessed for seaside and leeside stability 

following Melby and Kobayashi (2011). A description of these 

methodologies is presented in Appendix A. Three analyses for structure 

damage progression were performed to evaluate the effects of beach 

condition in front of the structure and storminess. First, a sensitivity 

analysis was performed varying elevations at the toe of the structure to 

represent the effect of varying beach conditions. Second, a damage 

progression analysis was performed for toe elevations obtained from the 

restoration measure model outputs described in section 2.2. Third, 

damage progression was performed for ten thousand 10-year life cycles 

without dividing the storms into storminess bins to corroborate 

storminess bin damage results. 

Some of the sampled storms produced water elevations that were above 

the crest elevation of the structure. In this condition, the seaside slope is 

no longer directly exposed to wave action. For these conditions, a cut-off 

elevation was implemented to the seaside damage equations to eliminate 

damage accumulations for when the water level hydrograph was 0.30 m

over the crest height.

4.3.1 Damage progression sensitivity analysis 

A sensitivity analysis of structure damage progression was performed to get 

a broad understanding of the effect of different beach conditions in front of 

the structure, storminess, and RSLC. Toe depths were gradually varied from 

the with-beach condition elevation to the no-beach design depth that were 

used for the peaks analysis in section 4.2. Seaside and leeside damage 

results for toe depths of -0.49, 0.5, 1.5, 2, and 2.44 m are shown in Tables 

4-5, 4-6, and 4-7 for each storminess bin and RSLC condition. Figures 4-2 

to 4-7 visually represent the damage variation surface of the data presented 

in the tables. The figures also include the no-damage upper limit surface 

(S = 5) in green and the initiation of failure damage surface (S = 20) in red 

for assessing the damage results for seaside stability and S=10 for leeside. 
ERDC/CHL TR-20-5 29

The damage progression results show that the presence of a beach in 

front of the structure reduces accumulated damage on both seaside and 

leeside for all storminess bins and SL cases. As expected, damage also 

increases with increases in RSLC. Note that although figures for seaside 

show a vertical scale of S of up to 100, once S is above levels 

corresponding to approximately S = 20, the value of S loses significance 

since it is above the range of physical modeling experimental testing on 

which it is typically defined. 

Table 4-5. Seaside and leeside damage for storminess bin 1. 

Bin 

Depth 

(m) 

RSLC 

(m) SSEA SLEE 

ST1 -0.49 0 0.0 0.0

ST1 -0.49 0.3 0.0 0.2

ST1 -0.49 0.5 0.0 0.4

ST1 -0.49 1.0 0.2 2.2

ST1 0.5 0 0.2 0.8

ST1 0.5 0.3 0.5 1.8

ST1 0.5 0.5 1.1 3.0

ST1 0.5 1.0 4.2 7.0

ST1 1.5 0 4.6 3.3

ST1 1.5 0.3 8.3 5.3

ST1 1.5 0.5 11.7 6.9

ST1 1.5 1.0 22.2 11.1

ST1 2 0 11.7 4.9

ST1 2 0.3 18.1 7.0

ST1 2 0.5 23.3 8.4

ST1 2 1.0 38.3 12.2

ST1 2.44 0 21.7 6.1

ST1 2.44 0.3 30.6 8.0

ST1 2.44 0.5 37.6 9.4

ST1 2.44 1.0 57.4 13.0
ERDC/CHL TR-20-5 30

Figure 4-2. Seaside damage variation surface for ST1. 

Figure 4-3. Leeside damage variation surface for ST1. 
ERDC/CHL TR-20-5 31

Table 4-6. Seaside and leeside damage 

for storminess bin 2.

Bin 

Depth 

(m) 

RSLC 

(m) SSEA SLEE 

ST2 -0.49 0 0.3 1.1

ST2 -0.49 0.3 0.5 1.8

ST2 -0.49 0.5 0.6 2.4

ST2 -0.49 1.0 1.2 4.4

ST2 0.5 0 2.1 2.6

ST2 0.5 0.3 3.0 4.0

ST2 0.5 0.5 3.9 5.3

ST2 0.5 1.0 7.8 9.4

ST2 1.5 0 11.3 5.6

ST2 1.5 0.3 16.3 7.9

ST2 1.5 0.5 20.8 9.7

ST2 1.5 1.0 34.8 14.1

ST2 2 0 23.2 7.4

ST2 2 0.3 32.1 9.8

ST2 2 0.5 39.5 11.6

ST2 2 1.0 61.1 15.7

ST2 2.44 0 40.1 8.9

ST2 2.44 0.3 53.3 11.3

ST2 2.44 0.5 64.0 13.0

ST2 2.44 1.0 93.9 16.8
ERDC/CHL TR-20-5 32

Figure 4-4. Seaside damage variation surface for ST2. 

Figure 4-5. Leeside damage variation surface for ST2. 
ERDC/CHL TR-20-5 33

Table 4-7. Seaside and leeside damage for storminess bin 3. 

Bin 

Depth 

(m) 

RSLC 

(m) SSEA SLEE 

ST3 -0.49 0 0.9 2.6

ST3 -0.49 0.3 0.9 3.5

ST3 -0.49 0.5 1.1 4.3

ST3 -0.49 1.0 1.6 6.8

ST3 0.5 0 4.4 4.5

ST3 0.5 0.3 5.3 6.1

ST3 0.5 0.5 6.2 7.6

ST3 0.5 1.0 10.6 12.6

ST3 1.5 0 19.3 8.0

ST3 1.5 0.3 24.7 10.7

ST3 1.5 0.5 29.8 12.8

ST3 1.5 1.0 46.3 18.2

ST3 2 0 36.9 10.2

ST3 2 0.3 46.9 13.2

ST3 2 0.5 55.6 15.3

ST3 2 1.0 81.7 20.3

ST3 2.44 0 61.5 12.1

ST3 2.44 0.3 76.9 15.0

ST3 2.44 0.5 89.6 17.1

ST3 2.44 1.0 126.1 21.8
ERDC/CHL TR-20-5 34

Figure 4-6. Seaside damage variation surface for ST3. 

Figure 4-7. Leeside damage variation surface for ST3. 

4.3.2 Incorporation of uncertainty

Uncertainty was considered in the structure damage progression analysis 

for the restoration measures. Two main types of uncertainties are 

generally recognized: aleatory and epistemic. Aelatory uncertainty relates 

to the natural randomness of a process while epistemic is associated with a 

lack of knowledge associated with models and data used to represent a 

system. The analysis incorporated uncertainty related to storm forcing, 

structure physical parameters, and empirical structure response model 

parameters. The statistical parameters for the sources of uncertainty 
ERDC/CHL TR-20-5 35

considered are shown in Table 4-8. Refer to Appendix A for parameter 

descriptions. The uncertainty associated with storm sampling was 

quantified through the implementation of the life-cycle approach. The 

total uncertainty was ultimately represented by confidence levels of 10%, 

16%, 84%, and 90%. 

Table 4-8. Table of statistical quantities where N 

(mean, standard deviation) denotes normally 

distributed parameter and D (value) represents 

a deterministic value. 

Parameter Distribution Description 

Hm0 from data from data

Tm from data from data

Nz from data from data

h from data from data

cotα N(p, 0.1) p = 2

P N(p, 0.05) p = 0.4

∆ N(p,0.06) p = 1.65

Ks D(0.4) for initial damage

Km1 N(5, 0.3) -

Km2 N(5, 0.4) -

Ks D(1) -

Kls1 N(125, 10) -

Kls2 N(0.4, 0.1) -

M50 N(450, 45) -

4.3.3 Damage progression for restoration measures

The damage progression results presented in section 4.3.1 covered a 

range of structure toe depths between the no-beach condition and the 

with-beach condition. It was shown that the variation of this parameter 

had a significant effect on the computed damage. To harmonize the 

Katrina Cut structure analysis with the Dauphin Island evolution 

modeling for the proposed restoration measures, the DEMs produced at 

the end of the 10-year modeling periods were used to extract toe 

elevations for the structure response modeling. Restoration measures R4 

through R7 impact the Katrina Cut area and were the focus of the 

analysis. The interpretation of the morphology analysis results was that 

the structure was considered fully exposed for storminess bin 3 for all 
ERDC/CHL TR-20-5 36

restoration measures except 7. Profiles were extracted from transects 

along the Katrina Cut, and the location of the structure toe was identified 

from the 2016 topographic survey. Model grid resolution turned out to be 

inadequate to resolve elevations at the structure toe; therefore, elevations 

from the breach near the west end of the structure for ST3SL3 were 

examined for use as toe elevations. The transect from which these 

elevations were extracted is shown as a red line in Figure 4-8. It is placed 

on the west side of the resulting barrier island breach. The ST3SL3 

profiles at the end of the 10-year simulation period for restoration 

measures R0, R4, R5, R6, and R7 are shown in Figure 4-9. The location 

at which the toe elevations were extracted is shown as a black dashed 

line. The figure also lists the toe elevations that represent each 

restoration measure.

Figure 4-8. Profile line for toe depth selection at known breaching point. 
ERDC/CHL TR-20-5 37

Figure 4-9. Profiles extracted from ST3SL 3 DEMs for R0, R4, R5, R6, and R7 

restoration measures at the end of 10-year simulation. 

The damage progression analysis was performed for the barrier island 

modeling output of restoration measures R4 through R7, represented by 

different toe elevations based on the ST3SL3 results. ST2SL1 was analyzed 

for the same toe elevations even though breaching did not occur for the 

restoration measure. This exercise was done to vary forcing conditions for 

the analysis and for comparison with the reliability computed from storm 

peaks. In practical terms, the morphology outputs for ST2SL1 show that 

the beach remains in front of the structure over the analyzed 10-year 

period for all restoration measures, and therefore the expected structure 

damage is small. The ST2SL1 resulting shorelines can be seen in section 

2.2, starting with Figure 2-9 for R0. 

The probability of failure for the damage progression analysis was 

evaluated by computing the percent of realizations per each storminess bin 

that the rubble mound structure breaches. As mentioned at the beginning 

of the section, damage values over S=20 are near the ultimate limit state 

and the structure would be near the initiation of breaching. 

The resulting mean damage values at the end of the 10-year simulations 

are shown in Table 4-9. The confidence levels associated with the 

computed damage values are presented in Tables 4-10 and 4-11 for seaside 
ERDC/CHL TR-20-5 38

and leeside damage, respectively. The results were consistent with the 

previous analysis with respect to the effect of the beach. Reliability 

estimates were determined for the damage progression assessment based 

on the percent of realizations that cause breaching at the end of the 

10-year simulation period. 

Table 4-9. Damage progression for STSL3 restoration measures representative 

toe depths. 

Case 

Restoration 

Measure 

Toe Depth 

(m) 

RSLC 

(m) Ssea 

Breach 

Percent 

(%) Slee 

Breach 

Percent 

(%) 

ST2SL1 R7 -0.39* 0.3 0.8 0.0 1.9 0.0

ST2SL1 R5 1.745* 0.3 25.4 11.0 9.6 1.0

ST2SL1 R6 1.825* 0.3 28.9 15.1 10.0 1.2

ST2SL1 R0 1.91* 0.3 32.5 18.5 10.2 1.8

ST2SL1 R4 2.283* 0.3 52.6 39.1 11.6 4.0

ST3SL3 R7 -0.39 1 1.9 0.0 6.7 0.0

ST3SL3 R5 1.745 1 68.2 49.7 19.9 23.4

ST3SL3 R6 1.825 1 74.6 56.9 20.3 23.7

ST3SL3 R0 1.91 1 79.6 59.9 20.9 26.3

ST3SL3 R4 2.283 1 123.9 78.0 22.8 28.6

* ST3SL3 toe depths.

Table 4-10. Accumulated mean seaside damage and confidence limits for restoration 

measures at the end of 10-year simulation period. 

Case 

Restoration 

Measure 

S 

(10%) 

S 

(16%) 

S 

(50%) 

S 

(84%) 

S 

(90%) 

ST2SL1 R7 0.0 0.0 1.0 1.4 1.6

ST2SL1 R5 14.2 16.0 25.4 35.2 37.6

ST2SL1 R6 14.6 16.5 28.2 39.7 46.1

ST2SL1 R0 18.3 21.4 30.9 46.1 52.2

ST2SL1 R4 28.4 30.7 47.9 73.7 82.3

ST3SL3 R7 1.1 1.2 1.8 2.7 3.2

ST3SL3 R5 42.1 48.2 65.1 90.9 100.5

ST3SL3 R6 49.5 53.1 70.8 95.4 101.9

ST3SL3 R0 51.6 57.2 74.1 105.0 108.3

ST3SL3 R4 81.1 90.4 114.9 158.4 170.9
ERDC/CHL TR-20-5 39

Table 4-11. Accumulated mean leeside damage and confidence limits for restoration 

measures at the end of 10-year simulation period. 

Case 

Restoration 

Measure 

S 

(10%) 

S 

(16%) 

S 

(50%) 

S 

(84%) 

S 

(90%) 

ST2SL1 R7 0.4 0.6 2.0 3.1 3.3

ST2SL1 R5 6.4 7.3 9.2 12.1 13.2

ST2SL1 R6 6.4 7.2 9.7 12.8 13.7

ST2SL1 R0 7.3 7.7 9.9 12.8 13.2

ST2SL1 R4 8.1 8.8 11.4 14.2 15.4

ST3SL3 R7 4.8 5.2 6.3 8.6 9.1

ST3SL3 R5 14.6 15.7 19.2 23.6 25.5

ST3SL3 R6 15.4 16.4 19.7 24.1 25.5

ST3SL3 R0 15.8 16.5 20.5 25.6 26.7

ST3SL3 R4 17.0 18.1 21.7 26.8 29.9

A comparison between reliability computed using peaks as described in 

section 4.2 and reliability estimates based on damage progression is 

presented in Table 4-12. For this comparison, the damage progression 

modeling was executed using the same -.49 m and 2.44 m toe depths used 

in the peaks analysis, which were representative of the with-beach and nobeach conditions. The results show comparable trends in terms of 

structure performance with respect to the previous analyses. The 

reliability for the no-beach condition was similar for ST2SL1 while for 

ST3SL3 the reliability based on damage progression was lower. For the 

with-beach condition, the reliability value was 1 for the damage 

progression stability modeling for both cases. The results reinforce the 

idea that the with-beach condition does not breach while the no-beach 

condition breaches.

Table 4-12. Comparison of computed reliability using 

peaks and storm hydrograph. 

Reliability - No Beach 

Case Peaks D. Progression 

ST2SL1 0.49 0.50

ST3SL3 0.33 0.19

Reliability – With Beach

ST2SL1 0.94 1

ST3SL3 0.79 1
ERDC/CHL TR-20-5 40

4.3.4 Damage progression analysis without storminess bins

Up to this point, the analyses have considered the decadal-scale variation 

in storm frequency and intensity through the classification of sampled 

10-year storm sequences into storminess bins. This classification was 

based on the barrier island morphologic response for each of the 1,000 

realizations performed for the analysis. It did not consider the coastal 

rubble mound structure response modeling. Because this structure was 

not designed to resist extreme events without catastrophic damage, 

analysis using a 10-year series was reasonable to define performance in the 

context of morphological behavior. It was of interest to compare damage 

computed under the prior life-cycle storminess approach with a sampling 

approach that did not involve storminess bins. The analysis was performed 

for 10,000 life cycles, without classifying each sampled 10-year realization 

into a storminess bin. In this approach, each realization maintains the 

storm recurrence rate. A large number of life cycles were sampled to 

ensure stability of the computed mean damage and evaluate the 

appropriateness of the number of realizations used for the storminess 

bins. This analysis would be expected to quantify the probability of 

damage over the full range of hazard that this structure will be exposed to 

over a 10-year period. 

The damage progression modeling for the 10,000 life cycles was 

performed for RSLC SL0, SL1, and SL3 using the response measure toe 

elevations from the ST3SL3 modeling. SL0 was added to evaluate the 

effect of the full range of RSLC on the structure response. An evaluation of 

ST1SL0 was included in the storminess bin analysis to have a basis of 

comparison with all three storminess bins. The seaside and leeside mean 

damage for both approaches is presented in Tables 4-13 and 4-14, 

respectively. The mean damage results for the 10,000 life-cycle approach 

closely match the mean damage results for storminess bin 2 for both the 

seaside and leeside. The 10,000 life-cycle mean damage is significantly 

higher than the damage of ST1, while the mean damage compared with 

ST3 is noticeably lower. These results reasonably confirm that the 

storminess bin 2 is representative of the decadal-scale variation in storm 

frequency and intensity corresponding to the uncensored TC climate.
ERDC/CHL TR-20-5 41

Table 4-13. Seaside damage comparison between no-storminess bin and storminess 

bin approaches. 

Case 

Restoration 

Measure 

Seaside S Breaching (%) 

No Bins With Bins Difference No Bins With Bins Difference 

ST1SL0 R7 0.6 0.0 0.6 0 0 0

ST1SL0 R5 19.7 9.8 9.9 4.8 0 4.8

ST1SL0 R6 22.2 10.4 11.7 6.5 0 6.5

ST1SL0 R0 24.7 11.3 13.3 8.8 0 8.8

ST1SL0 R4 39.7 21.2 18.6 23.5 4.7 18.8

ST2SL1 R7 0.9 0.8 0.1 0 0 0

ST2SL1 R5 27.8 25.4 2.4 11.5 11 0.5

ST2SL1 R6 30.7 28.9 1.8 14.9 15.1 0.2

ST2SL1 R0 34.3 32.5 1.7 18.7 18.5 0.2

ST2SL1 R4 53.6 52.6 1.0 38.9 39.1 0.2

ST3SL3 R7 1.4 1.9 0.5 0 0 0

ST3SL3 R5 52.2 68.2 16.0 43.5 49.7 6.2

ST3SL3 R6 56.9 74.6 17.7 48.5 56.9 8.4

ST3SL3 R0 62.3 79.6 17.4 53.7 59.9 6.2

ST3SL3 R4 91.2 123.9 32.7 70.7 78 7.3

Table 4-14 Leeside damage comparison between no-storminess bin and storminess 

bin approaches. 

Case 

Restoration 

Measure 

Leeside S Breaching (%) 

No Bins With Bins Difference No Bins With Bins Difference 

ST1SL0 R7 1.5 0.2 1.4 0 0 0

ST1SL0 R5 7.2 4.7 2.5 0.1 0 0.1

ST1SL0 R6 7.6 4.9 2.7 0.1 0 0.1

ST1SL0 R0 7.9 5.5 2.4 0.2 0 0.2

ST1SL0 R4 9.2 6.5 2.7 0.7 0 0.7

ST2SL1 R7 2.0 1.9 0.2 0 0 0

ST2SL1 R5 9.6 9.6 0.1 0.9 1 0.1

ST2SL1 R6 10.0 10.0 0.0 1.2 1.2 0

ST2SL1 R0 10.3 10.2 0.1 1.6 1.8 0.2

ST2SL1 R4 11.7 11.6 0.1 3.9 4 0.1

ST3SL3 R7 5.0 6.7 1.7 0 0 0
ERDC/CHL TR-20-5 42

Case 

Restoration 

Measure 

Leeside S Breaching (%) 

No Bins With Bins Difference No Bins With Bins Difference 

ST3SL3 R5 15.9 19.9 4.0 17.1 23.4 6.3

ST3SL3 R6 16.1 20.3 4.2 18.1 23.7 5.6

ST3SL3 R0 16.5 20.9 4.4 19.2 26.3 7.1

ST3SL3 R4 17.5 22.8 5.3 22.9 28.6 5.7
ERDC/CHL TR-20-5 43

5 Wave Transmission

The Katrina Cut rubble mound structure assessment included evaluating 

the transmitted wave conditions on the leeside due to wave overtopping of 

the structure. Wave transmission coefficients (Ct) and transmitted wave 

heights (Hm0_t) were computed for ST2SL1 and ST3SL3 scenarios used in 

the damage progression analysis. The impact of restoration measure 5 on 

transmitted wave heights was also assessed. 

5.1 Wave transmission analysis 

The wave transmission analysis was applied to all life cycles for each of the 

storminess bins taking into the account the variation of crest height, toe 

depth, and RSLC. The combinations of these three parameters each 

correspond to a specific configuration, and each configuration was 

evaluated separately for ST1 and ST3. Figure 5-1 shows a schematic 

diagram of the wave transmission analysis configurations. Although crest 

reduction was not modeled explicitly as part of the damage progression 

analysis, survey data show that the structure has been reduced at various 

locations. Two values of crest elevations of 0.91 m and 1.83 m were 

modeled in addition to the design crest height of 2.44 m for the wave 

transmission analysis. Wave transmission was evaluated using the

empirical relationship described in van der Meer et al. (2005). Refer to 

Appendix B for a description of this empirical relationship as well as its 

comparison to the wave transmission formula from the Coastal 

Engineering Manual (USACE 2002). 

The transmitted wave heights (Hm0_t) at the leeside of the structure were 

computed using the empirical relationships for the wave transmission 

coefficient, C_t, (refer to Appendix B) and the incident wave heights at the 

toe of the structure obtained from the Monte Carlo simulation (Hm0_i). The 

C_t was applied to all the configurations mentioned previously, and the 

maximum of each life cycle was averaged to get a representative Hm0_t for 

the specific configuration. Table 5-1 shows the maximum mean 

transmitted wave height for all storminess bins, toe depths, and SLR 

configurations with a crest height of 2.44 m. The results for the full 

analysis are presented in Appendix C. As expected, wave transmission 

increases with storminess bin and sea level rise. This is due to the 

structure being submerged for longer periods of time, hence dissipating 

less wave energy. The same is true when evaluating lower crest heights 
ERDC/CHL TR-20-5 44

than the one shown in Table 5-1. The computed maximum mean 

transmitted wave heights are used in the following subsection and 

Appendix C to evaluate the impact restoration measure R5 would have on 

transmitted waves. 

Figure 5-1. Schematic diagram of wave transmission analysis configurations. 

Table 5-1. Mean maximum transmitted waves for all storminess 

bins, toe depths, and RSLC configurations with a crest height 

of 2.44 m. 

Bin 

Toe 

Depths 

(m) 

Hm0_t (m) 

SL0 SL1 SL2 SL3 

1 -0.39 0.02 0.04 0.04 0.15

1 1.75 0.10 0.11 0.12 0.21

1 1.83 0.11 0.13 0.16 0.49

1 1.91 0.11 0.12 0.13 0.28

1 2.28 0.14 0.28 0.43 0.79

2 -0.39 0.15 0.31 0.43 0.75

2 1.75 1.15 1.39 1.55 1.89

2 1.83 0.27 0.49 0.64 1.00

2 1.91 0.94 1.18 1.33 1.67

2 2.28 0.37 0.61 0.76 1.12

3 -0.39 1.75 1.88 1.97 2.17

3 1.75 0.87 1.12 1.29 1.65

3 1.83 1.48 1.71 1.85 2.14

3 1.91 0.75 0.99 1.16 1.53

3 2.28 1.91 2.11 2.22 2.42
ERDC/CHL TR-20-5 45

5.2 Impact of restoration measure 5 on transmitted wave energy

Restoration measure 5 is implemented in the leeside of the Katrina Cut 

structure. Figure 5-2 shows with a red polygon the footprint of the areas 

with an elevation difference between the proposed tidal flats and marsh 

habitat restoration and the existing conditions. This measure has the 

potential to impact the wave transmission to the leeside of the structure as 

it significantly reduces the water depths. The profile along the length of 

the structure was extracted for existing conditions and for R5 morphologic 

modeling results. These profiles include the initial condition for R5 and 

the end of simulation period profiles for ST2SL1 and ST3SL3. The location 

of the profile line is shown in Figure 5-2 as a yellow line 30 m from the 

structure crest, and the profiles are shown in Figure 5-3. 

Restoration measure 5 increases the bottom elevation of the area at the 

back of the Katrina Cut structure to an elevation of near 0.5 m. As seen in 

Figure 5-3, the existing profile in the lee of the structure is variable and 

has depths over 2 m at the deepest point. This location is identified in plan 

view in Figure 5-2. The initial R5 profile is shown as the black dashed line 

in Figure 5-3 while the morphology modeling result for ST2SL1 is 

represented by the blue line. The bottom elevation is reduced less than 

0.30 m at the end of the 10-year simulation period throughout the profile. 

For ST3LS3, the elevation profile continuously decreases towards the west 

end of the Katrina Cut where breaching is predicted to occur. 

An assessment of the impact of restoration measure 5 on wave 

transmission was performed by evaluating the transmitted wave heights 

with respect to the resulting water depths. The wave breaking depth 

limitation criteria discussed in section 3.2 was used to determine if the 

mean maximum transmitted waves were susceptible to wave breaking. The 

occurrence of wave breaking on the leeside of the Katrina cut structure for 

transmitted waves is shown in Figures 5-4 to 5-6 for the initial R5 

condition and three crest height elevations and five toe depths. Conditions 

where wave breaking occurs are identified with a red circle, and conditions 

where the transmitted waves are not depth limited are identified with a 

green square. These figures are meant to give a general idea of the effect of 

R5 on transmitted waves at the deepest location in the existing conditions 

profile. Breaker criteria analysis for the other configurations are shown in 

Appendix C. Since water depths vary along the profile at the end of the 

10-year simulations, in particular for ST3SL3, profiles of maximum 

transmitted wave heights (Hm0_max) were generated that take into account 
ERDC/CHL TR-20-5 46

depth limitation. Figures C-28, C-29, and C-30 from Appendix C show the 

maximum transmitted wave heights along the length of the profile leeside 

of the Katrina Cut structure for three crest height elevations. 

Figure 5-2. Line for profile extraction of water depths and footprint of restoration 

measure 5 on the leeside of the Katrina Cut structure. 

Figure 5-3. Comparison between restoration measure 5 profiles and existing profile 

(NAVD88). 

Wave transmission 

depths location
ERDC/CHL TR-20-5 47

Figure 5-4. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and SLR configurations with a crest height of 2.44 m. 

Leeside depth: -0.47 m. 

Figure 5-5. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and SLR configurations with a crest height of 2.44 m. 

Leeside depth: -0.47 m. 
ERDC/CHL TR-20-5 48

Figure 5-6. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and SLR configurations with a crest height of 2.44 m. 

Leeside depth: -0.47 m. 
ERDC/CHL TR-20-5 49

6 Summary and Conclusions

The rubble mound structure placed at the Katrina Cut in Dauphin Island, 

AL, was designed to be a temporary structure whose ultimate objective 

was to prevent oil migration into the Mississippi Sound after the Deep 

Water Horizon oil spill. As such, the armor stone was undersized with 

respect to generally accepted armor design criteria that limits catastrophic 

damage from extreme storm events. Within the first 2 years of its 

construction, the structure showed damage as the seaside stone armor was 

directly exposed to wave action. The survey performed in 2016 revealed 

areas with significant damage and lowering of the crest height at multiple 

locations. Since construction, morphological changes near the structure 

have been characterized by the steady accumulation of sand and the 

formation of a broad beach on the seaside (south) of the rubble mound 

structure. The assessment of the Katrina Cut rubble mound structure was 

performed as part of the Alabama Barrier Island Restoration Assessment 

study undertaken to evaluate potential restoration measures for Dauphin 

Island considering varying RSLC and storminess scenarios.

The rubble mound structure response was evaluated using a life-cycle 

modeling approach. The analysis considered the outcomes of the 

morphologic modeling of the barrier island performed by the USGS for 

proposed restoration measures that impact the Katrina Cut area. For the 

morphological response analysis, one thousand 10-year life cycles were 

sampled and classified based on storm frequency and intensity into three 

storminess bins. Morphological modeling of measures was performed for 

an intermediate scenario of storminess bin 2 (ST2) and low RSLC scenario 

(SL1) and an extreme scenario of ST3 and high RSLC scenario (SL3). 

The beach in front of the structure plays an important role in protecting 

the structure. The beach dissipates wave energy through wave breaking 

due to depth limitation. An initial assessment of the effect of the beach on 

the reliability of the structure showed that the absence of the beach can 

reduce its reliability more than 50%. The structure is expected to be 

protected for scenario ST2SL1 for all restoration measures since the beach 

was retained. Scenario ST3SL3 results in a loss of beach for restoration 

measures R4, R5, and R6 and corresponding expected breaching 78, 50, 

and 57 percent of the time for seaside damage, respectively. No breaching 

is expected for restoration alternative R7 since a beach is retained seaward 
ERDC/CHL TR-20-5 50

of the structure. If the structure holds for ST3SL3, breaching occurs to the 

west of the structure at the end of the 10-year period. 

Since the Katrina Cut structure is low crested, overtopping wave 

transmission is expected. Wave transmission analysis was applied for all 

life cycles for each of the storminess bins taking into the account the 

variation of crest height, toe depth, and sea level rise. Mean maximum 

transmitted wave heights were computed for the various configurations 

and, as expected, wave transmission increases with storminess bin and 

with sea level rise. Restoration alternative 5 is implemented in the leeside 

of the Katrina Cut structure and increases sea floor elevations there. The 

effect of R5 is to reduce transmitted wave heights compared to existing 

conditions. Wave transmission is increased at the end of the 10-year 

simulation period for ST3SL3 due to a decrease in the profile, but it still 

remains higher than the existing conditions profile. 
ERDC/CHL TR-20-5 51

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Survey Open-File Report 2020-1001. Reston, VA: US Geological Survey.
ERDC/CHL TR-20-5 52

Nadal‐Caraballo, N. C., J. A. Melby, V. M. Gonzalez, and A. T. Cox. 2015. North Atlantic 

Coast Comprehensive Study – Coastal Storm Hazards from Virginia to Maine. 

ERDC/CHL TR-15-5. Vicksburg, MS: US Army Engineer Research and 

Development Center.

Nadal-Caraballo, N. C., V. M. Gonzalez, and L. Chouinard. 2019. Storm Recurrence Rate 

Models for Tropical Cyclones. ERDC/CHL TR-19-4. Vicksburg, MS: US Army 

Engineer Research and Development Center.

Resio, D. T., S. J. Boc, L. Borgman, V. Cardone, A. T. Cox, W. R. Dally, R. G. Dean, D. 

Divoky, E. Hirsh, J. L. Irish, D. Levinson, A. Niedoroda, M. D. Powell, J. J. 

Ratcliff, V. Stutts, J. Suhada, G. R. Toro, and P. J. Vickery. 2007. White Paper on 

Estimating Hurricane Inundation Probabilities. Consulting Report prepared by 

USACE for FEMA. Vicksburg, MS: US Army Engineer Research and 

Development Center.

Roelvink, D., A. Reniers, A. van Dongeren, J. van Thiel de Vries, R. McCall, and J. 

Lescinski. 2009. "Modelling Storm Impacts on Beaches, Dunes and Barrier 

Islands." Coastal Engineering 56(11-12): 1133–1152

Smith, J. M., D. T. Resio, and A. K. Zundel. 1999. STWAVE: Steady-State Spectral Wave 

Model. Report 1 User’s Manual for STWAVE Version 2.0. CHL-99-1. Vicksburg, 

MS: US Army Corps of Engineers Waterways Experiment Station. 

Sommers, P. 2003. A Poisson Model for Hurricanes of the North Atlantic. Middlebury 

College Economics Discussion Paper No. 03-24. Middlebury College, Department 

of Economics.

SWAN Team 2014. SWAN Scientific and Technical Documentation SWAN Cycle III, 

Version 41.01. Delft University of Technology, The Netherlands. 

http://swanmodel.sourceforge.net/

URS/Dewberry. 2012. Joint Probability Analysis of Hurricane Flood Hazards for Gulf, 

Bay, Walton, Okaloosa, Santa Rosa, and Escambia Counties Florida, and 

Baldwin and Mobile Counties, Alabama. Federal Emergency Management 

Agency (FEMA) Map Modernization program for the Northwest Florida Water 

Management District. Boulder, CO.

USGS (US Geological Survey), State of Alabama, National Fish and Wildlife Foundation, 

and USACE Mobile District. 2017. Alabama Barrier Island Restoration 

Assessment. Interim Report. 

Van Gent, M. R. A., and B. Pozueta. 2005. “Rear-Side Stability of Rubble Mound 

Structures.” Coastal Engineering 2004, 3481–93. World Scientific Publishing 

Company. 

van der Meer, J. W., R. Briganti, B. Zanuttigh, and B. Wang. 2005. “Wave Transmission 

and Reflection at Low-Crested Structures: Design Formulae, Oblique Wave 

Attack and Spectral Change.” Coastal Engineering 52(10–11): 915–929.

van der Meer, J. W. and K. d’Angremond. 1991. “Wave Transmission at Low Crested 

Structures.” Proceedings of the Coastal Structures and Breakwaters Conference, 

Institution of Civil Engineers. London UK: Thomas Telford Publishing. 
ERDC/CHL TR-20-5 53

Westerink, J. J., R. A. Luettich Jr., A. M. Baptista, N. W. Scheffner, and P. Farrar. 1992. 

“Tide and Storm Surge Predictions Using Finite Element Model.” Journal of 

Hydraulic Engineering 118(10): 1373–1390. 

US Army Corps of Engineers (USACE). 2002. Coastal Engineering Manual. EM 1110-2-

1100. Washington, DC: US Army Corps of Engineers. 
ERDC/CHL TR-20-5 54

Appendix A: Stability and Damage Analysis

This appendix discusses the stability equations used for the Katrina Cut 

structure. 

A.1 Seaside armor stability

The seaside armor stability was based on maximum momentum flux 

methods of Melby and Hughes (2004) with generalizations from Melby 

and Kobayashi (2011). The approach captures the effects of wave steepness

and water depth on stability and damage progression over a wide range 

water depths primarily because these processes are included in the wave 

momentum flux basic principles included in derivation of the equations. 

The maximum wave momentum flux is highly nonlinear for nonlinear 

waves (i.e., steep waves in shallow water). This corresponds to the case 

where armor stability is at its minimum. The nonlinear wave momentum 

flux was approximated numerically using a Fourier solution. The resulting 

approximate relation was found to be 

max

.

.

.

.

A

F

w m

m

m

M h A

ρ gh gT

H A

h

H A

h





          

    

 

   

    

 

   

1

2 2 0

2 026

0

0

0 391

0

1

0 639

0 180

(A-1)

where Hm0 is the significant wave height, h is the local water depth, Tm is 

the mean wave period, Mf is the momentum flux, ρw is the density of 

water, and g is the acceleration of gravity. A nonlinear approximation for 

momentum flux is important because stability is at its minimum when the 

incident wave is the most nonlinear.

Two stability equations resulted from the fit to data using Equation A-1. 

The recommended equations for stability are
ERDC/CHL TR-20-5 55

.

m

m s z

S N

a K N

      

 

    

0 2

1

(A-2)

and

/

max (/ )

( )

F w

m

r n

M γ h h N

S D

      

      

1 2 2

1 50

(A-3)

Plunging waves:

. cot m m mc

m

a s s

K P θ   0 18

1

1

(A-4)

Surging waves:

/

. ./ (cot )

P

m

m P P m mc

m m

s

a s s

K P θ s    

3

0 18 0 5 3

2

(A-5)

, / . cot . mc m m mc s HL s  0  0 0035 0 028 θ  (A-6)

Solving for the stable stone size yields

/ /

max (/ )

Δ

F w

n m

s z

S M γ h D ha

K N

                

        

1 5 1 2 2

50 (A-7)

Here, Dn50 = (V50)1/3 = nominal stone size, V50 = M50/ρr = median volume 

of armor stone, M50 = median mass, cot θ = structure seaward slope, ∆ = Sr

– 1, Sr = ρr/ρw = relative specific gravity of armor stone, ρr = density of 

stone, ρw = density of water that the structure is in, Nm is the momentum 

flux stability number, P = notional permeability of the structure, S = 

Ae/(Dn50)2 = normalized eroded area, Ae = eroded area, and Nz = storm 

duration/Tm. S and Ae are illustrated in Figure A-1. Ks = 1.3 is an empirical

parameter to account for accelerated damage that occurs with constant 

wave conditions, γw is the specific weight of water, θ is the seaside 

structure slope, and sm is the local wave steepness. Km1 and Km2 are 

empirical fit coefficients.
ERDC/CHL TR-20-5 56

Figure A-1. Illustration of rubble mound 

damage parameters. 

Damage progression was assessed for seaside and leeside stability 

following Melby and Kobayashi (2011) and the methodology presented in 

Melby et al. (2015). The equations for damage progression used for the 

assessment of Katrina Cut structure are presented in the latter. 

A.2 Leeside armor stability

Leeside stability was determined using Van Gent and Pozueta (2004) 

stability equations, revised to be in similar form to the seaside equations 

(Melby 2009). 

/

% ,

Δ

r

ls m

n ls

ls z

S u T

D a

K N



      

                    

1

1 10

50 125

(A-8)

. / / (cot ) [ exp( / )] r r als φ R H c rear s

    2 5 1 1 10 (A-9)

where Sls is the leeside damage, r = 6 for constant wave conditions, Tm-1,0 = 

m-1/m0 of incident spectrum, Tm-1,0 ~ Tp /1.1 for a JONSWAP incident 

wave spectrum, Rc-rear = freeboard of leeside edge of crest, cot φ = leeside 

slope, u1% = maximum crest velocity exceeded by 1% of the waves, Hs = 

Hm0 of incident wave spectrum, Kls and r are empirical fit parameters, and 

Dn ls ( ) and ∆ls are the nominal stone size and density parameter for the 
ERDC/CHL TR-20-5 57

leeside armor, respectively. A leeside stability number, Nls, is introduced 

where

/

% ,

( Δ)

r

m ls

ls

n ls ls ls z

u T S N

D KN a

        

                 

1

1 10 1

125

(A-10)

Then, the single storm leeside damage for constant wave conditions can be 

expressed as

 

r

S K N aN ls ls z ls ls  (A-11)

The nth moment of the incident wave energy density spectrum is given by

( ) n m f S f df n



 

0

(A-12)

The crest velocity exceeded by 1% of the waves is given by 

.

. %

%

.( )

.

c

f C

f s

s c

s

z R

γ

u γ H

gH B

H



      

 

            

     

0 5

0 5 1

1

1 7

1 01

(A-13)

where γf-C = friction factor on crest, γf = friction factor on seaward slope, Rc

= freeboard of seaside crest, Bc = breakwater crest width, and z1% = runup 

exceeded by 1% of waves. The friction coefficients and runup can be

computed by Equations 14 and 15, respectively:

,

, ,

,

.

. *( ) .

.

s

f fC s s

s

ξ

γ γ ξ ξ

ξ



 



  

   

    



  

1

1 1

1

0 55 2

0 05625 2 0 55 2 10

1 0 10

(A-14)

% , ,

, , /

s s

s s s

z c ξ for ξ p

γH c c ξ for ξ p

 

 

   





  



1 0 1 1

12 1 1

(A-15)
ERDC/CHL TR-20-5 58

Here c2 = 0.25 c1

2/c0, p = 0.5 c1/c0, and γ = γf = reduction factor for 

roughness and angular wave attack. The Iribarren parameter based on the 

first negative moment wave period is 

,

,

tan s

s

m

θ ξ

H

L





1 

1 0

(A-16)

where

,

,

m

m

gT L

π



 

2

1 0

1 0 2 (A-17)

and c0 = 1.45 and c1 = 5.1 for z1%. For the Katrina Cut rubble mound 

structure, Equation A-15 becomes

% , ,

, ,

. .

. ./ .

s s

s s s

z ξ for ξ

H ξ for ξ

 

 

   



   

1 1 1

1 1

1 45 1 76

5 10 4 485 1 76 (A-18)

A.3 Reliability

Reliability is defined in terms of a limit state that characterizes the failure 

surface of the system. The limit state can be formulated in terms of 

performance, gg, which is obtained by comparing the strength (R) of the 

system against the load (S) on the system:

gg = RR − SS (A-19)

The probability of failure is given by 

PPff = PP(gg ≤ 0) (A-20)

and the reliability is given by

RRff=1-PPff (A-21)

Armor seaside and leeside stability and damage equations can be 

rearranged into resistance (strength) and forcing (load) terms. The limit 
ERDC/CHL TR-20-5 59

state equation for seaside armor stability was obtained by rearranging 

Equation A-7 as

gg = √∆DDnnSS0.2

aamm

− hNNzz

0.1 �

(MMFF)mmmmmm

ρρwwggh2 �

0.5

(A-22)

Equation A-8 was rearranged to obtain the leeside damage limit state 

equation, expressed as

gg = √∆DDnnnn SS1/6

aa − NNzz

1/12 �

uu1%TTmm−1,0

KK � (A-23)
ERDC/CHL TR-20-5 60

Appendix B: Wave transmission

The formulas for wave transmission for rock armored low crested 

structures by van der Meer and D'Angremond (1991) and van der Meer et 

al. (2005) were evaluated for application in current study. Both equations 

consist of empirical relationships that estimate the wave transmission 

coefficient (Ct), which is defined as

CCtt = HHmm0_tt

HHmm0_ii

(B-1)

where HHmm0_ii is the incident significant wave height and HHmm0_tt the 

transmitted wave height. 

B.1 Comparison of wave transmission formulas

The Coastal Engineering Manual (USACE 2002) Table VI-5-15 provides a 

wave transmission coefficient (Ct) equation for conventional low-crested 

structures developed from physical model experiments (van der Meer and 

D'Angremond 1991):

CCtt = �0.031 HHss

DDnn50

− 0.24� RRcc

DDnn50

+ bb (B-2)

bb = −5.42 SSoooo + 0.0323 HHss

DDnn50

− 0.0017 ( BBcc

DDnn50

)1.84 + 0.51 (B-3)

SSoooo = HHss

LLoooo

(B-4)

RRcc = vvcccccc − ɳ (B-5)

where Hs = incident significant wave height at structure toe, ɳ = 

deepwater water level, Sop = deep water wave steepness based on the peak 

period, Lop = deepwater wavelength based on the peak period, Dn50 = 

seaside nominal stone size, Rc = structure freeboard, cccccc = structure 

crest elevation. The parameter b varies depending on the type of 

structure being analyzed; for this study, Dauphin Island’s structure was 

modeled as a conventional structure. The results obtained from Equation 

B-2 are bounded to 0.075 ≤ CCtt ≤ 0.75 where the expected behavior is that 

Ct will increase as the structure becomes more submerged (Rc<0) until it 

reaches the upper bound.
ERDC/CHL TR-20-5 61

A second formulation for computing wave transmission by van der Meer et 

al. (2005) was considered. Based on Equation B-2, it was developed using 

an expanded database that included a broader range of rubble mound 

designs and forcing results. Wave transmission was estimated as

CCtt = �

−0.4 RRcc

HHss

+ 0.64 � BB

HHss

�

−0.31

∗ (1 − eeee(−0.5 ξξoooo)),

B

Hs

< 10

−0.35 RRcc

HHss

+ 0.51 � BB

HHss

�

−0.65

∗ (1 − eeee(−0.41 ξξoooo)),

B

Hs

> 10

(B-6)

ξξoooo = tan(θθ)

SSoooo

(B-7)

where ξξoooo= Irribaren parameter based on the peak period and tan(θθ) = is 

the inverse of the structure’s seaside slope. Similar to Equation B-2, the 

results obtained by this new relationship are bounded to 0.075 ≤ Ct ≤ Kut

were the upper bound limit is defined by Equation B-8.

KKuu = −0.006 BB

HHss

+ 0.93 (B-8)

Applying equation for a conventional structure which truncates values at 

0.075 ≤ Ct ≤ Kut where Kut is defined by

KKuu = −0.006 BB

HHss

+ 0.93 (B-9)

Finally, with the estimated wave transmission coefficient, the transmitted 

wave height by overtopping at the leeside of the structure was obtained 

applying rearranging equation B-1:

HHmm0_tt = HHmm0_ii ∗ CCtt (B-10)

B.2 Assessment of wave transmission equations

Wave transmission coefficients were computed for all life cycles and timesteps using equations B-2 and B-6. The latter was selected for the wave 

transmission analysis based on a comparison of the results from the two 

equations. Figure B-1 shows an example application of Equation B-2. It 

can be seen that for values of Rc/Dn50 greater than -4, Equation B-2 

produces a reduction in Ct as the structure becomes more submerged. In 

contrast, Figure B-2 shows that Equation B-6 computed Ct values keep 
ERDC/CHL TR-20-5 62

increasing to its maximum as the structure becomes more submerged. 

Therefore, Equation B-6 was selected to compute wave transmission 

values in this study.

Figure B-1. Wave transmission coefficient for SB 2 using Equation B-2. 

Figure B-2. Wave transmission coefficient for SB 2 using Equation B-6. 
ERDC/CHL TR-20-5 63

Appendix C: Wave Transmission Analysis 

Results (Figures)

Figure C-1. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: -0.47 m. 

Figure C-2. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 1.83 m. 

Leeside depth: -0.47 m. 
ERDC/CHL TR-20-5 64

Figure C-3. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 2.44 m. 

 Leeside depth: -0.47 m. 

Figure C-4. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: -0.47 m. 
ERDC/CHL TR-20-5 65

Figure C-5. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 1.83 m. 

Leeside depth: -0.47 m. 

Figure C-6. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: -0.47 m. 
ERDC/CHL TR-20-5 66

Figure C-7. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 0.91 m. 

 Leeside depth: -0.47 m. 

Figure C-8. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 1.83 m. 

 Leeside depth: -0.47 m. 
ERDC/CHL TR-20-5 67

Figure C-9. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: -0.47 m. 

Figure C-10. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: -0.19 m. 
ERDC/CHL TR-20-5 68

Figure C-11. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 1.83 m. 

 Leeside depth: -0.19 m. 

Figure C-12. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: -0.19 m. 
ERDC/CHL TR-20-5 69

Figure C-13. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: -0.19 m. 

Figure C-14. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 1.83 m. 

Leeside depth: -0.19 m. 
ERDC/CHL TR-20-5 70

Figure C-15. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: -0.19 m. 

Figure C-16. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: -0.19 m. 
ERDC/CHL TR-20-5 71

Figure C-17. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 1.83 m. 

Leeside depth: -0.19 m. 

Figure C-18. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: -0.19 m. 
ERDC/CHL TR-20-5 72

Figure C-19. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: 0.38 m. 

Figure C-20. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 1.83 m. 

Leeside depth: 0.38 m. 
ERDC/CHL TR-20-5 73

Figure C-21. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST1 toe depths and RSLC configurations with a crest height of 2.44 m. 

 Leeside depth: 0.38 m. 

Figure C-22. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: 0.38 m. 
ERDC/CHL TR-20-5 74

Figure C-23. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 1.83 m. 

 Leeside depth: 0.38 m. 

Figure C-24. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST2 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: 0.38 m. 
ERDC/CHL TR-20-5 75

Figure C-25. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 0.91 m. 

Leeside depth: 0.38 m. 

Figure C-26. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 1.83 m. 

Leeside depth: 0.38 m. 
ERDC/CHL TR-20-5 76

Figure C-27. Mean maximum transmitted waves and wave breaking classification at 

leeside for all ST3 toe depths and RSLC configurations with a crest height of 2.44 m. 

Leeside depth: 0.38 m. 

Previous figures provide the mean maximum transmitted waves in the 

leeside of the structure at location near the center of the structure. 

Nevertheless, the longitudinal east-west profile along the back of the 

structure at the end of the 10-year simulation period for restoration 

measure 5 slopes westward at the end of ST3SL3. The profiles of 

maximum transmitted wave heights (Hm0_t) based on ST2SL1 and ST3SL3 

forcing and the effects of depth limitation considering their respective 

profiles at the end of the 10-year simulation period are shown in 

Figures C-28, C-29, and C-30 for structure crest heights of 0.91 m, 1.81 m, 

and 2.44 m. 
ERDC/CHL TR-20-5 77

Figure C-28. Restoration measure 5 modeling conditions and leeside bottom 

influence on transmitted waves for 0.91 m crest height. 

Figure C-29. Restoration measure 5 modeling conditions (after 10 years) and leeside 

bottom influence on transmitted waves for 1.81 m crest height. 
ERDC/CHL TR-20-5 78

Figure C-30. Restoration measure 5 modeling conditions and leeside bottom 

influence on transmitted waves for 2.44 m crest height. 
ERDC/CHL TR-20-5 79

Acronyms and Abbreviations

1D one-dimensional

ADCIRC Advanced Circulation 

CRL coastal reference location 

DEM digital elevation model 

FEMA Federal Emergency Management Agency

JPM-OS joint probability method with optimal sampling 

NAVD88 North American Vertical Datum of 1988

CA1D one-dimensional model framework

RSLC relative sea level change 

SLC sea level change 

SRR storm recurrence rate

ST storminess bin

SWAN Simulating WAves Nearshore 

TC tropical cyclone 

USACE US Army Corps of Engineers

USGS US Geological Survey 
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PLEASE DO NOT RETURN YOUR FORM TO THE ABOVE ADDRESS.

1. REPORT DATE 

April 2020

2. REPORT TYPE

Final Report

3. DATES COVERED (From - To)

4. TITLE AND SUBTITLE

Alabama Barrier Island Restoration Assessment Life-Cycle Structure Response Modeling

5a. CONTRACT NUMBER

5b. GRANT NUMBER

5c. PROGRAM ELEMENT NUMBER

6. AUTHOR(S)

Victor M. Gonzalez, Fabian A. Garcia-Moreno, Jeffrey A. Melby, Norberto C. NadalCaraballo, and Elizabeth S. Godsey

5d. PROJECT NUMBER

452808

5e. TASK NUMBER

5f. WORK UNIT NUMBER

7. PERFORMING ORGANIZATION NAME(S) AND 

ADDRESS(ES)

Coastal and Hydraulics Laboratory

US Army Engineer Research and 

Development Center

3909 Halls Ferry Road

Vicksburg, MS 39180-6199

Mobile District

US Army Corps of Engineers

109 St Joseph St.

Mobile, AL 36602

8. PERFORMING

ORGANIZATION REPORT

NUMBER

ERDC/CHL TR-20-5

9. SPONSORING/MONITORING AGENCY NAME(S) AND ADDRESS(ES)

Mobile District, US Army Corps of Engineers

Mobile, AL 36602

10. SPONSOR/MONITOR'S ACRONYM(S)

SAM USACE

11. SPONSOR/MONITOR'S

REPORT NUMBER(S)

12. DISTRIBUTION/AVAILABILITY STATEMENT

Approved for public release; distribution is unlimited.

13. SUPPLEMENTARY NOTES

14. ABSTRACT

Dauphin Island, a barrier island off the coast of Alabama, plays an important role in the protection of the state’s coastal natural resources. 

In 2011, the State of Alabama constructed a rubble mound berm across a 2 km breach in the western end of the barrier island to prevent 

oil from the Deep Water Horizon oil spill from migrating into the Mississippi Sound. The breach, referred to as the Katrina Cut, was 

caused by Hurricane Katrina in 2005. The US Army Corps of Engineers and the US Geological Survey (USGS) performed the Alabama 

Barrier Island Restoration Assessment study to assess the current and future function of Dauphin Island and evaluate potential restoration 

measures. A Monte Carlo life-cycle structure response assessment of the Katrina Cut rubble mound structure was performed as part of 

this study by the US Army Engineer Research and Development Center, Coastal and Hydraulics Laboratory. Damage, wave 

transmission, and reliability were computed within the context of the decadal barrier evolution analysis performed by the USGS for 

various storminess and relative sea level change scenarios. The presence of a beach in front of the structure plays an important role in its 

protection. The breaching potential for measures was evaluated.

15. SUBJECT TERMS 

Barrier islands, Dauphin Island (Ala.), Ocean waves, Rubble mound breakwaters, Shorelines—Monitoring, Storms

16. SECURITY CLASSIFICATION OF: 17. LIMITATION OF 

ABSTRACT

SAR

18. NUMBER 

OF 

PAGES

90

19a. NAME OF RESPONSIBLE PERSON

Victor M. Gonzalez a. REPORT

Unclassified

b. ABSTRACT

Unclassified

c. THIS PAGE

Unclassified 19b. TELEPHONE NUMBER (Include area code)

601-634-3006

Standard Form 298 (Rev. 8/98)

Prescribed by ANSI Std. Z39.18