Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-1-1 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Appendix 1 

Terminology 

Event tree analysis is an inductive analysis that depicts the sequence of occurrences that 

shows the logical sequence of the occurrence of events in, or states of, a system following an 

initiating event. 

A failure mode is a way that failure can occur, described by the means by which element or 

component failures must occur to cause loss of the sub-system or system function. 

Fault tree analysis is a systems engineering method for representing the logical combinations 

of various component states and possible causes that can result in a specific system state (called 

the top event). 

A fragility curve is a function that defines the probability of failure, conditioned on some 

appropriately defined intensity such as an applied load, a velocity, flood elevation, or other 

parameter. 

A hazard is a threat, which may result from either an external cause (e.g. earthquake, flood, 

or human agency) or an internal vulnerability, with the potential to initiate a failure mode. It is a 

source of potential harm or a situation with a potential to cause loss. 

The performance of a system or component is its ability to meet functional requirements. The 

performance of an item was described by various elements, such as flood protection, reliability, 

capability, efficiency, and maintainability. The design and operation of the system affects this 

performance. 

A system is an entity comprising an interacting collection of discrete elements and commonly 

defined using deterministic models. The word deterministic implies that the system is identifiable and not uncertain in its architecture. The definition of the system is based on analyzing its 

functional and/or performance requirements. A description of a system may be a combination of 

functional and physical elements. A system was divided into subsystems that interact. Additional 

details in the definition of the system lead to a description of the physical elements, components, 

and various aspects of the system. Methods to address uncertainty in systems architecture are 

available and were employed as provided by Ayyub and Klir (1996). 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Reliability is the ability of a system or a component to fulfill its design functions under 

designated operating and/or environmental conditions for a specified time period. This ability is 

commonly measured using probabilities. Reliability is, therefore, the probability that the failure 

event, however defined, does not occur. 

Consequences are damages or losses from some failure event. Each failure of a system has 

some consequence(s). A failure could cause economic damage, environmental damage, injury or 

loss of human life, or other possible events. Consequences need to be quantified in terms of 

failure-consequence severities using relative or absolute measures for various consequence types 

to facilitate risk analysis. 

Risk is the potential of losses for a system resulting from an uncertain exposure to a hazard or 

as a result of an uncertain event. Risk should be based on identified risk events or event 

scenarios. Risk is a multi-dimensional quantity that includes event-occurrence probability, eventoccurrence consequences, consequence significance, and the exposed population; however, it is 

commonly measured as a pair of the probability of occurrence of an event, and the outcomes or 

consequences associated with the event’s occurrence. Another common representation of risk is 

in the form of a curve depicting specified losses and the probability of exceeding those losses. 

Probability is a measure of the likelihood, chance, odds, or degree of belief that a particular 

outcome will occur. A conditional probability is the probability of event occurrence based on the 

assumption that another event (or multiple events) has occurred. 

Safety was defined as the judgment of risk tolerance (or acceptability in the case of decision 

making) for the system. Safety is a relative term since the decision of risk acceptance may vary 

depending on the individual or the group of people making the judgment. 

Risk analysis is the technical and scientific process to breakdown risk into its underlying 

components. Risk analysis provides the processes for identifying hazards, event-probability 

assessment, and consequence assessment. The risk analysis process answers three basic 

questions: (1) What can go wrong? (2) What is the likelihood that it will go wrong? (3) What are 

the consequences if it does go wrong? Also, risk analysis can include the impact of making any 

changes to a system to control risks. 

Risk Assessment an examining of the tradeoffs that must take in any effort directed toward 

risk mitigation or risk reduction. 

Risk communication was defined as an interactive process of exchange of information and 

opinion among stakeholders such as individuals, groups, and institutions. It often involves multiple messages about the nature of risk or expressing concerns, opinions, or reactions to risk 

managers or to legal and institutional arrangements for risk management. Risk communication 

greatly affects risk acceptance and defines the acceptance criteria for safety. 

A scenario is a unique combination of circumstances that lead to an outcome of interest. 

Thus there may be loading scenarios, failure scenarios or downstream flooding scenarios. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-1 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Appendix 2 

New Orleans East Basin 

NOE – Background 

The New Orleans East hurricane protection system was designed as part of the Lake 

Pontchartrain, LA, and Vicinity Hurricane Protection Project. The New Orleans East (NOE) 

portion of the project protects 45,000 acres of urban, industrial, commercial, and ecological 

lands. As designed, the levees were generally constructed with a 10-foot crown width with side 

slopes of 1 on 3. The height of the levees varies but was in the range of 12–19 feet, depending 

upon location. There are also various types of floodwall segments along the line of protection. 

As designed, there is a total of approximately 206,000 linear feet of levees and floodwalls, 

eight pump stations, three U.S. Fish and Wildlife Service (USFWS) pump stations, a multitude 

of culverts through/over the levee/floodwall, and multiple gate closures for road and rail 

crossings. The NOE basin is essentially broken into two major sections, as shown in Figure NOE 

1. The west side of the basin is primarily residential and the east side is essentially a wetlands 

area. These two areas are separated by a small levee. The west side of the basin is further divided 

into residential and industrial areas. The area along the GIWW and IHNC is primarily industrial 

while the remainder of the western portion is residential in nature. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-2 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 1. New Orleans East Basin – Major Stretches by DM 

NOE – Design Memorandums 

For the purposes of the IPET Task 10 risk assessment, each basin must be broken into 

“reaches” that are defined by a combination of physical characteristics, major elevation changes, 

and potential consequences. Many of the basic reaches were defined by when individual design 

memorandums (DM) were completed and then constructed since different stretches of the 

levee/floodwall were raised at different times throughout the life of the NOE protection system. 

There are a total of 7 levee/floodwall major stretches separated by different DM’s within NOE. 

These 7 are defined below and illustrated in Figure 1. 

• Lakefront Airport Floodwall 

o Beginning Point: Northwest corner of basin below Ted Hickey Bridge 

o Ending Point: End of floodwall just south of Hayne Blvd closure gate 

• Citrus Lakefront Levee/Floodwall 

o Beginning Point: Begin transition levee just south of Hayne Blvd closure 

o Ending Point: Levee height transition at Paris Road and USFWS levee 

• Lakefront Levee 

o Beginning Point: Levee transition at Paris Road and USFWS interior levee 

GIWW 

MRGO 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

o Ending Point: South Point at northeast end of basin 

• East Levee 

o Beginning Point: South Point at northeast corner of basin 

o Ending Point: GIWW at southeast corner of basin 

• East Back Levee 

o Beginning Point: GIWW at southeast corner of basin 

o Ending Point: Northeast end of Michoud Canal floodwall 

• Citrus Back Levee/Floodwall 

o Beginning Point: Northeast end of Michoud Canal floodwall 

o Ending Point: Southwest corner of basin at IHNC 

• IHNC East Levee/Floodwall 

o Beginning Point: Southwest corner of basin at IHNC 

o Ending Point: Northwest corner of basin under Ted Hickey Bridge 

NOE Basin – Layout of Reaches for Risk Model (Pre-Katrina) 

Within these major stretches defined by the DM’s there are reaches, which are defined by 

physical changes in the protection system, i.e. switching from floodwall to levee, etc..., or by 

changes in geotechnical parameters. Within each reach, there are specific “key points” whose 

reliability needs to be determined in order to calculate the effect on the overall reach being 

evaluated. An example of a “key point” would be a closure gate at a road or rail line crossing 

along a floodwall. In addition, there are transition points between walls and levees that are also a 

critical part of the risk analysis, particularly since several of these transition points were 

locations where significant scour damage occurred during Katrina. Task 10 engineers reviewed 

existing plans, damage survey reports, and conducted field verification inspections to ensure 

each basin was accurately defined within the system. As a part of the field verification 

inspections, GPS coordinates were obtained and stationing from DM’s and “as-built” plans were 

verified. For each basin, this information was transformed into a spread sheet and then a system 

map for each basin, as shown in Figure NOE 2 and further clarified in Figure NOE 3 for 

“congested” areas where several shorter reaches are close together. Finally, digital photographs 

with incorporated notes were developed to compliment the spread sheets and system map for 

further clarification. This collection of information was then categorized to get a clear picture of 

how the basin should be defined for risk assessment purposes. There are a couple of interior, 

local levees, but these are not considered substantial hurricane protection systems and are only 

used to define the interior drainage within the basin itself. Therefore, they are not shown and 

defined within the context of the risk model other than for flow characteristics. 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 2. New Orleans East Basin – Reaches Defined 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 3. New Orleans East Basin – Insets 1 &2 (Refer to Figure NOE 2) 

Task 10 basin definition starts at the northwest corner of the basin where the floodwall along 

the IHNC intersects the floodwall along the Lakefront Airport (Reach NOE1). This occurs at 

Sta. 4+02 B/L, which is equal to the DM stationing of 10+13 W/L. The end of the physical 

definition of the NOE basin occurs at the same point since it is self enclosed. Refer to 

Figures NOE 2 and NOE 3 for the general location of the reach as the basic characteristics are 

detailed within this narrative. 

Reach NOE1 (Lakefront Airport DM) 

This reach is defined by 2,326 linear feet of concrete I-wall along the Lakefront Airport. It is 

located at the northwest end of the basin. There are two key points (NOE1a and NOE1b) within 

this reach, both closure gates, located near the end of this reach. Relatively short T-wall sections 

surround the gate closures. The reach ends just after the second closure gate for Hayne 

Boulevard. The weighted average elevation of the top of this wall was approximately elevation 

11.6 (NAVD88 2004.65) based upon a physical survey taken November 2005. It had an average 

free standing height of approximately 8.2 feet based upon field measurements following Katrina. 

There was significant scour from limited overtopping and/or wave splash along this section of 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

I-wall, as shown in Figure NOE 4, but there was no permanent deformation of the wall. Refer to 

the “Post Katrina NOE System Definition” section of this narrative for a detailed description of 

changes to this reach made by Task Force Guardian (TFG). 

Figure NOE 4. Scour Behind Lakefront Airport FW from Overtopping 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Reach NOE2 (Citrus Lakefront DM) 

This reach is defined by a short 97-ft transition levee between the end of the Lakefront 

Airport floodwall and the beginning of the west Stars & Stripes floodwall. There are no key 

points within this reach; however, there are two transition features where the levee ties into the 

concrete capped I-walls on both ends of the reach. One of these transition points (between Reach 

NOE 1 and NOE 2) is shown in Figure NOE 5 where the scour occurred during Katrina. The 

weighted average height of this reach was 13.3 (NAVD88 2004.65) based upon an adjusted 

LIDAR done prior to Katrina, although at the transition between NOE1 and NOE2 the reach is 

significantly lower until the railroad embankment is encountered. 

Reach NOE3 (Citrus Lakefront DM) 

This reach is defined by 2,325 linear feet of concrete capped I-wall. This reach is commonly 

referred to as the west Stars & Stripes floodwall. There are two basic types of floodwall along 

this reach each consisting of about 1⁄2 the length of this reach. The first type is a concrete capped 

I-wall with levee on both sides and the second is a concrete I-wall section where the protected 

side has a concrete sidewalk adjacent to a road. The risk assessment model focused on both of 

these with an emphasis on the I-wall section without the concrete sidewalk which helps to serve 

as erosion protection when the wall is overtopped. There are no key or transition points within 

this reach. The weighted average top of wall elevation along this reach is elevation 13.5 

(NAVD88 2004.65) based upon a physical survey taken in March 2006. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-8 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 5. Transition Point Between Reaches (NOE 1 & NOE 2) (Viewed from Reach NOE2 looking 

north towards Lakefront Airport) 

Reach NOE4 (Citrus Lakefront DM) 

This reach is defined by 2,330 linear feet of the Stars and Stripes levee. It is located between 

the west and east Stars and Stripes floodwalls. A small concrete I-wall for the discharge pipes at 

the St. Charles Pump Station is located near the east end of this reach. There are no key points 

within this reach, but there are two transition points where the levee abuts both Stars and Stripes 

floodwalls. The ends of the St. Charles Pump Station floodwall are not considered transition 

points since they are essentially flush with the top of the levee. The weighted average top 

elevation of this reach was 13.3 (NAVD88 2004.65) based upon an adjusted LIDAR survey 

taken prior to Katrina. The transition between the levee and the east Stars & Strips floodwall is 

shown in Figure NOE 6. 

Reach NOE5 (Citrus Lakefront DM) 

This reach is defined by 2,270 linear feet of concrete I-wall. It is commonly referred to as the 

east Stars and Stripes floodwall. There are no key or transition points within this reach. The 

weighted average top of wall elevation for this reach is 13.7 (NAVD88 2004.65) based upon a 

physical survey done in November 2005. This wall was not damaged during Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-9 

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Figure NOE 6. Transition between Reaches NOE4 (Levee) and NOE5 (FW) 

Reach NOE6 (Citrus Lakefront DM) 

This reach is defined by a 19,112 linear feet segment of levee. It begins at the end of the east 

Stars & Stripes floodwall and ends at the west side of the Lincoln Beach floodwall. There are a 

couple of short, small floodwall sections within this reach located at the Citrus and Jahncke 

Pump Stations; however, these are very short walls that are not considered significant within the 

overall characteristics of the reach. There are three transition points assigned to this reach. These 

are for the levee/concrete I-wall interface points where the east Stars & Stripes floodwall ends, 

the Jahncke Pump Station floodwall, and where the levee ties into the west side of the Lincoln 

Beach floodwall. The floodwall at the Citrus Pump Station was not considered as a transition 

point since the top of the wall was essentially flush with the top of the levee. The weighted 

average elevation for the top of this reach was 12.9 (NAVD88 2004.65) based upon an adjusted 

LIDAR survey taken prior to Katrina. There was some minor overtopping and/or wave splash 

over this levee at various locations, as indicated in Figure NOE 7, but no significant damage. 

Reach NOE7 (Citrus Lakefront DM) 

This reach is defined by a 1,474 linear feet of concrete I-wall located at Lincoln Beach. 

There is one “key point” located in the flood wall, which is a closure gate that is approximately 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-10 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

100 feet wide. Short concrete T-walls surround the gate. The weighted average top of wall 

elevation for this reach is 12.1 (NAVD88 2004.65) based upon a physical survey taken in 

November 2005. The wall was overtopped during Katrina and performed very well. The Lincoln 

Beach floodwall is shown in Figure NOE 8 for reference. 

Figure NOE 7. Minor Scour from Overtopping at Jahncke Pump Station 
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Figure NOE 8. Lincoln Beach Floodwall (Reach NOE7) (Looking west from Reach NOE8) 

Reach NOE8 (Citrus Lakefront DM) 

This reach of levee, 2,724 linear feet, ends the Citrus Lakefront section at the intersection of 

Paris Road, the interior local levee, and the west side of the Lakefront Levee. There are no key 

points within this reach, although the levee height is considerably different as it proceeds to the 

Lakefront Levee section, as shown in Figure NOE 9. The weighted average top elevation of this 

levee was 12.6 (NAVD 88 2004.65) based upon an adjusted LIDAR survey taken prior to 

Katrina. This section was overtopped and damaged during Katrina, primarily in the area where 

the lower Citrus Lakefront levee section transitions to the higher East Lakefront Levee section as 

shown in Figure NOE 9. This represents the only transition point within this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-12 

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Figure NOE 9. Begin Lakefront Levee at Citrus Lakefront and Paris Road (Lakefront Levee @ El. 18.6 +/- 

and Citrus Lakefront Levee @ 12.6+/-) 

Reach NOE9 (Lakefront Levee DM) 

This reach covers 33,165 feet of levee along Lake Pontchatrain from Paris Road to South 

Point, which is the extreme northeast corner of the basin. There is 368-ft-long I-wall around the 

Exxon/Mobil pipeline crossing that provides the only transition point within the reach. The 

weighted average top elevation of this reach was 18.6 (NAVD88 2004.65) based upon an 

adjusted LIDAR survey taken before Katrina. This reach was not overtopped or damaged to any 

significant extent during the storm. 

Reach NOE10 (East Levee DM) 

This reach is defined by a 27,665-linear-ft segment of levee from South Point to where 

Highway 90 crosses the levee. There are four transition points within this stretch, including three 

different drainage culvert headwalls on the levee toe. The other remaining transition point is for 

the floodwall surrounding the gated closure at Highway 11. The gated closure at Highway 11 

represents the only “key point” within the reach as well. The weighted average top elevation of 

this levee is 15.1 (NAVD88 2004.65) based upon an adjusted LIDAR survey done prior to 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Katrina. This section was slightly overtopped or subject to wave splash during Katrina, but 

suffered minimal damage other than at transition points. 

Reach NOE11 (East Levee DM) 

This levee is 8,942’ long and extends southeast from Highway 11 and serves as the section 

where the levee design was modified to account for wave action. There are no “key points” 

located within this reach. The weighted average top elevation of this reach was 16.7 (NAVD88 

2004.65) based upon an adjusted LIDAR survey taken prior to Katrina. It is believed that this 

section of levee suffered minor overtopping and/or wave splash, but overall it performed well. 

Reach NOE12 (East Levee DM) 

The final reach of levee along the East section is 7,190 ft long and extends to the GIWW 

with a weighted average top elevation of 15.0 (NAVD88 2004.65) based upon an adjusted 

LIDAR survey taken prior to Katrina. There are two transition points within this reach including 

a short sheet pile wall around a drainage structure and transition concrete I-walls around a 

railroad crossing. There is one key point, the railroad gated closure, which is included in this 

reach. At the time of Katrina, the railroad closure gate and transition walls were roughly 4 feet 

lower than the adjacent levee and was overtopped and heavily damaged during the storm. An 

aerial view of this damage is shown in Figure NOE 10. 
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Figure NOE 10. Aerial View of Damage at RR Closure Along East Levee 

Reach NOE13 (East Back Levee DM) 

This section of levee, measuring 22,257 linear feet, was heavily damaged during Katrina 

from overtopping. Prior to Katrina, it had a weighted average top elevation of 15.5 (NAVD88 

2004.65) based upon an adjusted LIDAR survey that was taken before the storm. It begins at the 

east end where it ties into the southern edge of the East Levee and continues to the east end of 

the floodwall around the Orleans Parish Pump Station #15. There are no key points within this 

reach. There is one transition point assigned to this reach for the floodwall/levee interface at the 

very western edge of this reach where it ties into the Pump Station #15 floodwall. Much of this 

levee was destroyed, as shown in Figure NOE 11, and was rebuilt by TFG. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-15 

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Figure NOE 11. Failure of Levee by Overtopping East of Pump Station #15 (East Back Levee) 

Reach NOE14 (East Back DM) 

This reach is defined by the floodwall around Orleans Parish Pump Station #15. There were 

two types of walls within this reach prior to Katrina, two 120-ft transition sheet pile walls at 

elevation 17.5 (NAVD88 2004.65) at the both ends of a middle 253-ft T-wall section with a top 

elevation of 22.2 (NAVD88 2004.65). The total length of this reach is 493 feet. Portions of the 

transition sheet pile sections were heavily damaged during Katrina as shown in Figure NOE 12. 

There are no key or transition points assigned to this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-16 

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Figure NOE 12. Floodwall Failure Near Orleans Pump Station #15 

Reach NOE15 (East Back DM) 

This 10,120-ft section of levee extends from the west end of the Orleans Parish Pump Station 

#15 floodwall to the start of the floodwall on the east side of the Michoud Canal at the GIWW. 

This levee section had a weighted average top elevation of 16.8 (NAVD88 2004.65) at the time 

of Katrina based upon an adjusted LIDAR survey taken prior to the storm. There are no key 

points, but two transition points assigned to this reach. Both transition points are for 

levee/floodwall interfaces at both ends of this reach. This reach was heavily damaged during 

Katrina and was rebuilt under TFG. 

Reach NOE16 (East Back DM) 

This reach consists of the east floodwall around the Michoud Canal. It is primarily concrete 

capped I-wall, but has a short transition sheet pile wall at the beginning of the reach. It is 

approximately 10,757 feet long and had a weighted average top elevation of 17.9 (NAVD88 

2004.65) based upon a physical survey taken in November 2005. The reach starts at the GIWW 

and continues along the Michoud Canal where it joins with the Citrus Back floodwall. There are 

18 key points along this reach for gated closures at industry and road crossings. However, from 

site inspections, it appears as if five of these gates are placed in the permanently closed position. 
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As shown in Figure NOE 13, the transition sheet pile floodwall at the beginning of this reach 

failed during Katrina; this section was rebuilt under TFG. The concrete capped I-wall section 

was overtopped and suffered significant scour as shown in Figure NOE 14, but performed well. 

There are no transition points assigned to this reach. 

Figure NOE 13. Floodwall Failure at East End of Michoud Canal FW 
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Figure NOE 14. Scour Damage Behind Michoud Canal Floodwall 
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Reach NOE17 (Citrus Back DM) 

The beginning of this reach starts at the northwest end of the Michoud Canal and ends at the 

southwest side of the Michoud Canal at the GIWW. This reach consists of 9,318 feet of 

floodwall with an average weighted top elevation of 20.8 (NAVD88 2004.65) based upon a 

physical survey taken in November 2005. This section was overtopped during Katrina but 

suffered only minor scour problems. There are no key or transition points assigned to this reach. 

Reach NOE18 (Citrus Back DM) 

This reach represents the 7,905 ft segment of levee between the Michoud Canal and Michoud 

Slip. There are no key points within this reach of levee, but there are two transition points where 

the levee ties into floodwalls on both ends. This reach had a weighted average top elevation of 

17.2 (NAVD88 2004.65) at the time of Katrina based upon an adjusted LIDAR survey taken 

prior to the storm. 

Reach NOE19 (Citrus Back DM) 

The reach represents the 6,155 ft of floodwall around the Michoud Slip. There are two key 

points (gated closure) within this reach, but no transition points assigned to this reach. This reach 

has a weighted average top elevation of 16.7 (NAVD88 2004.65) based upon a physical survey 

taken in November 2005. This section of wall suffered minimal scour damage during Katrina. 

Reach NOE20 (Citrus Back DM) 

This reach contains 15,940 feet of levee between the west end of the Michoud Slip and the 

east end of the combination floodwall for the bulk loading facility. There are two transition 

points assigned to this reach where the levee ties into floodwalls at both ends. There are no key 

points within this reach. This reach has a weighted average top elevation of elevation 14.0 

(NAVD88 2004.65) based upon an adjusted LIDAR survey taken before the storm. Portions of 

this levee were overtopped during Katrina with moderate areas of scour damage, as shown in 

Figure NOE 15, but there were no major breaches in this reach. 
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Figure NOE 15. Scour Damage Along Citrus Back Levee 

Reach NOE21 (Citrus Back DM) 

This reach is defined by the 1,820-ft combination floodwall built for the Bulk Loading 

Facility and Elaine Pump Station. This wall was heavily damaged during Katrina, as shown in 

Figure NOE 16, and was repaired under TFG. Prior to Katrina, it had a top elevation of 12.1 

(NAVD88 2004.65) based upon both physical survey and LIDAR information. There is one key 

point, gate N1 for the Bulk Loading Facility, within this reach. There are no transition points 

assigned to this reach. 
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Figure NOE 16. Floodwall Failure at Bulk Loading Facility/Elaine PS 

Reach NOE22 (Citrus Back DM) 

This reach is for the levee (3,453 ft long) between the floodwall at the Bulk Loading 

Facility/Elaine PS (east side of reach) and the floodwall that is just east of the Amid Pump 

Station (west side of reach). There are no key points within this reach, but there are two 

transition points where the levee ties into the adjacent floodwalls on both ends. It had a weighted 

average top elevation of 13.4 (NAVD88 2004.65) prior to Katrina based upon an adjusted 

LIDAR survey taken before the storm. There was minor overtopping of this reach during Katrina 

with no significant damage. 

Reach NOE23 (Citrus Back DM) 

This reach is the 1,587 ft section of floodwall located just east of the Amid Pump Station. 

This wall did suffer minor overtopping, but no major damage. It had a weighted average top 

elevation of 14.5 (NAVD88 2004.65) based upon a physical survey taken in November 2005. 

There are no key or transition points assigned to this reach. 
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Reach NOE24 (Citrus Back DM) 

The final reach of the Citrus Back Levee stretch is 2,348 feet of levee extending from the end 

of the floodwall just east of the Amid Pump Station to its tie in with the Inner Harbor Navigation 

Canal (IHNC) east levee. There is one key point (railroad closure gate) located within this reach 

and two transition points (railroad closure transition I-walls and the end of the I-wall near the 

Amid Pump Station). The transition around the railroad closure structure was overtopped and 

sustained significant scour damage, as shown in Figure NOE 17. This reach had a weighted 

average top elevation of 13.8 (NAVD88 2004.65) based upon an adjusted LIDAR survey taken 

prior to Katrina. 

Figure NOE 17. Erosion Damage Around RR Closure (Near GIWW and IHNC Intersection) 

Reach NOE25 (IHNC DM) 

This reach is 3,803 feet long and consists primarily of levee with several gated closures. The 

weighted average top elevation of this reach was 12.2 (NAVD88 2004.65) based upon an 

adjusted LIDAR survey taken prior to Katrina. There are four closure gates, each a key point, 

within this reach; all of which suffered erosion damage from overtopping during Katrina. Each 

of these four closure structures has short adjoining I-walls which represent four transition points 

that are assigned to this reach. There is one additional transition point assigned to this reach 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-23 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

where the end of the levee ties into the beginning of the floodwall, as shown in Figure NOE 19. 

A typical example of the scour damage around these closure structures is shown in Figure NOE 

18. Structural damage was minimal to these closure structures. The very end of this reach 

suffered a major washout area where the levee serves as a ramp just near the I-10 overpass. A 

photograph of this washout damage is shown in Figure NOE 19. 

Reach NOE26 (IHNC DM) 

This short reach of floodwall (537 ft) starts near the end of the washout area and extends just 

under the I-10 overpass. This section is considered a reach because it faces several different 

directions and contains two key points, both closure gates. There are no transition points 

assigned to this reach. The weighted average top elevation of this wall was 12.4 (NAVD88 

2004.65) based upon a physical survey of the wall completed in November 2005. This section 

was overtopped but received minimal scour damage behind the wall. 

Reach NOE27 (IHNC DM) 

This reach consists of a short transition levee (526 ft) between floodwalls. There are no key 

points within this short reach, but there are two transitions assigned to this reach where the 

floodwall interfaces with the levee. The weighted average top elevation of the levee was 12.6 

(NAVD88 2004.65) based upon a physical survey taken of this reach in November 2005. This 

reach was overtopped during Katrina, but suffered no significant damage. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-24 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 18. Scour Damage around Gated Closure Along IHNC 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-25 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 19. Major Washout Area from Overtopping Near I-10 Overpass (Note: Located Along IHNC 

East Side) 

Reach NOE28 (IHNC DM) 

This section of floodwall (1,876 ft) starts between the I-10 and Highway 90 overpasses and 

ends where it serves as the foundation for the Dupuy Storage Facility (see Figure NOE 20). 

There are no key points, but there is one transition point in this section, which is the old 

Highway 90 overpass location. It does not appear as if remedial repairs were made to this 

transition section when the overpass was relocated. The weighted average top elevation of this 

reach is 12.9 (NAVD88 2004.65) based upon a physical survey taken during November 2005. 

This reach was overtopped with minimal damage during Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-26 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 20. Floodwall Serves as Building Foundation (Dupuy Storage Facility – IHNC East) 

Reach NOE29 (IHNC DM) 

This short section of floodwall (643 ft) serves as the Dupuy Storage Building foundation as 

shown in Figure NOE 20. This section was deemed an individual reach because overtopping 

issues along this short reach warranted its treatment as a separate structure. The weighted 

average top elevation is 13.5 (NAVD88 2004.65) based upon a physical survey taken during 

November 2005.

Reach NOE30 (IHNC DM) 

The last reach of the basin consists of 8,168 ft of floodwall. There are several key points 

(closure gates) within this reach. There is one transition within this reach at the Jourdan Road 

crossing. The weighted average top elevation of this reach is 12.2 (NAVD88 2004.65) based 

upon a physical survey taken in November 2005. This includes the shorter sections of T-walls 

located around closure gates. Portions of this wall were overtopped as indicated by the erosion 

behind the floodwall adjacent to closure gate E-13 and shown in Figure NOE 21. This erosion, 

which measures approximately 8-ft wide by 2.5-ft deep, did not cause major structural problems 

for the wall at this location. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-27 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 21. Erosion Behind Floodwall Adjacent to Gate E-13 (IHNC East) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-28 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

NOE – Post-Katrina Layout of Reaches for Risk Model 

Reach NOE1 (Lakefront Airport DM) 

The only significant change made by TFG to this reach was to repair the scour damage along 

the wall and placement of a concrete scour pad on the protected side of the wall shown in 

Figure NOE 22. Refer back to Figure NOE 4 for a photograph of the wall following Katrina 

without the scour pad, which was taken from the same general location as Figure NOE 22, as 

noted by the road ramp on the right hand side of both photographs.

Figure NOE 22. Scour Protection Pad Behind Lakefront Airport FW 

Reach NOE2 (Citrus Lakefront DM) 

The only significant change made to this reach was improved scour protection around the 

transition points at both ends where the levees tie into floodwalls. An example of this improved 

protection is shown in Figure NOE 23, where the transition between Reach NOE 1 (floodwall) 

and NOE 2 (levee) has been raised and protected with a concrete pad. Figure NOE 23 is taken 

from the same general location as Figure NOE 5 which shows the transition before the 

improvement. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-29 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 23. Transition Improvements Between Reaches NOE 1&2 (Note: refer back to Figure NOE 5 

taken from same location) 

Reach NOE3 (Citrus Lakefront DM) 

The major improvement to this reach was the installation of a concrete scour pad on the 

protected side along the entire reach. There were no changes made to free standing wall heights 

in this reach from the pre-Katrina condition. 

Reach NOE4 (Citrus Lakefront DM) 

The only significant improvement to this reach, known as the Stars & Strips Levee, made by 

TFG was improved scour protection around the transitions located at both the east and west ends 

of this reach. Figure NOE 24 depicts the improved scour protection at the interface of Reaches 

NOE4 (Stars & Stripes Levee) and NOE5 (East Stars & Stripes Floodwall). Refer back to Figure 

NOE 6 for a photograph of this transition taken after Katrina but prior to TFG repairs. 

Reach NOE5 (Citrus Lakefront DM) 

There have been two major improvements to this reach. The first is the free standing height 

of this I-wall has been reduced from approximately 7.5 feet (pre-Katrina) to 6.0 feet (TFG 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-30 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

improvements). Secondly, a concrete scour protection pad has been added on the protected side 

of the wall for its entire length. 

Reach NOE6 (Citrus Lakefront DM) 

The only significant improvements to this reach were scour protection that was placed 

around all three transition points. An example of the scour protection around the Jahncke Pump 

Station Floodwall is shown in Figure NOE 25 at the same general location where the damage 

photo (Figure NOE 7) was taken following Katrina. 

Figure NOE 24. Transition Improvements Between Reaches NOE 4&5 (Note: refer back to Figure NOE 6 

taken from same location) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-31 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 25. Transition Improvements at Jahncke Pump Station (Note: refer back to Figure NOE 7 

taken from same location) 

Reach NOE7 (Citrus Lakefront DM) 

The major improvements to the Lincoln Beach floodwall were a reduction in the free 

standing height of the I-wall sections and scour protection. The free standing height has been 

reduced from approximately 10 feet (pre-Katrina) to 5.5 feet (post Katrina repairs under TFG). A 

concrete scour protection pad was added at the base of the wall on the protected side. A view of 

the improvements (under construction) to the Lincoln Beach floodwall is shown in 

Figure NOE 26, taken from roughly the same vantage point as Figure NOE 8, which is at the east 

edge of the wall looking west from Reach NOE8. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-32 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 26. Lincoln Beach Floodwall Repairs (Note: refer back to Figure NOE 8 taken from same 

location) 

Reach NOE8 (Citrus Lakefront DM) 

The only improvement made to the last reach associated with this DM was improved scour 

protection around the east end of the Lincoln Beach floodwall transition into this reach. This 

transition point is assigned to Reach NOE8 and scour protection at this interface, as shown in 

Figure NOE 26, was provided by TFG. The transition between the end of this DM and the 

beginning of the higher Lakefront Levee section has not been significantly changed from preKatrina conditions. 

Reach NOE9 (Lakefront Levee DM) 

This reach was not damaged during Katrina and no major changes were made under TFG 

with the exception of transition point scour improvements around the Exxon/Mobil floodwall 

pipeline crossing. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-33 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Reach NOE10 (East Levee DM) 

This reach suffered minimal damage during Katrina and there are no major changes to the 

characteristics of this reach itself. However, scour protection improvements were made to three 

of the four transition points within this reach. These repairs were made at the drainage structure 

headwalls, as typically shown in Figure NOE 27. The other transition point, the floodwall at the 

Highway 11 gated closure, had no changes made to it under TFG. 

Reach NOE11 (East Levee DM) 

There were no improvements made to this reach or any of the transitions within this reach. 

Figure NOE 27. Scour Repairs to Drainage Structure Headwalls (Note: Reach NOE10 drainage structure 

repairs – typical) 

Reach NOE12 (East Levee DM) 

There were significant repairs made to select areas of this reach that were heavily damaged 

during Katrina. The first major repair was the entire levee was rebuilt from the CSX railroad 

closure structure southward to where the levee ties into the East Back Levee at the GIWW. 

Therefore, the pre-Katrina weighted average elevation of 15.0 was restored under TFG 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-34 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

(NAVD88 2004.65). The second major improvement was the reconstruction of the CSX railroad 

closure gates and adjoining floodwalls. At the time Katrina hit, this gated closure and supporting 

floodwall structure was roughly 4 feet lower than the adjoining levees and it was heavily 

damaged, as shown in Figure NOE 10. A new gate and supporting floodwall was installed, as 

shown in Figure NOE 28. Note that the new gate and supporting walls were constructed such 

that it is no longer 4 feet lower than the adjacent levees. 

Figure NOE 28. Reconstruction of CSX Railroad Closure Structure 

Reach NOE13 (East Back Levee DM) 

This section of levee suffered heavy damage from overtopping during Katrina, and it was 

totally rebuilt under TFG. The new levee was constructed to elevation 19.5 (NAVD88 2004.65) 

with a final design grade of elevation 17.5 to account for 2 feet of long-term settlement. Thus, 

this reach will be approximately 4 feet higher than the pre-Katina condition as constructed and 

2 feet higher once the long-term settlement occurs, and it reaches its design elevation. Not only 

was the levee raised but it was also built with better materials and with improved methods. Prior 

to the Katrina, this reach had been constructed with hydraulic fill taken from the MRGO/GIWW, 

which included a mix of sand and silt materials. Within the context of this analysis, the preKatrina soil properties for this reach were classified as “high to very highly erodible” in terms of 

resistance to overtopping damage. Since the new levee was constructed with better material and 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-35 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

better methods (not hydraulically filled), the erodibility classification was improved to the 

“moderate” category. Figure NOE 29 shows the newly constructed levee. It can be compared 

back to the failed levee along this reach shown in Figure NOE 11. 

Figure NOE 29. New Levee Along the GIWW (East Back Levee) 

Reach NOE14 (East Back Levee DM) 

This 493-ft reach of combination floodwall was also heavily damaged during Katrina and 

was the focus of major repairs by TFG. The surrounding sheet pile I-walls failed during Katrina 

(see Figure NOE 12) and were replaced primarily with a concrete T-wall at elevation 22.2 

(NAVD88 2004.65). There is a very short (18 feet) transition of shorter concrete I-wall at 

elevation 20.2 on the west end of the repaired section. Thus, the weighted average elevation of 

this overall reach has been raised from 19.9 (pre-Katrina) to 22.1 (post-Katrina repairs under 

TFG). The newly repaired wall is shown in Figure NOE 30. 

Reach NOE15 (East Back Levee DM) 

This reach of levee failed by overtopping during Katrina and was rebuilt by TFG. The new 

levee was built to elevation 19.5 with a final design grade of elevation 17.5 (NAVD88 2004.65) 

to account for 2 feet of long-term settlement. The pre-Katrina elevation of this reach was 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-36 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

approximately 16.8. Thus, as constructed, the new levee is roughly 2.7 feet higher with a final 

height 0.7 feet higher once it settles to its final grade. Like the Reach NOE13 repairs, this new 

levee was also constructed with more erosion resistant material and with improved construction 

methods when compared to the pre-Katrina condition. 

Figure NOE 30. Repaired Floodwalls Around Pump Station #15 

Reach NOE16 (East Back Levee DM) 

Many parts of this reach of floodwall were significantly overtopped during Katrina, 

including a failed section of transition sheet pile near the beginning of the reach, as shown in 

Figure NOE 13. This short section of transition sheet pile I-wall has been replaced with a 

concrete I-wall by TFG. In addition, the free standing height of the non-failed concrete I-wall 

sections was reduced from approximately 7.5 feet (pre-Katrina) to 4.5 feet (post-Katrina). Also, 

a concrete scour pad, as shown in Figure NOE 31, has been provided on the protected side of the 

concrete I-wall sections. Both the shortened free standing height and scour protection pad are 

evident in the figure. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-37 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 31. Michoud Canal Floodwall Improvements 

Reach NOE17 (Citrus Back Levee DM) 

Consistent with the east side of the Michoud Canal, the west side also had its free standing 

height reduced and scour protection provided on the back side. During a visual inspection held 

during June 2006, TFG construction crews were making the noted improvements as evident in 

Figure NOE 32. The free standing height was reduced to 7.0 feet (post-Katrina) from 9.3 feet 

(pre-Katrina). A scour protection pad was provided on the back side. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-38 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 32. West Floodwall Repairs around Michoud Canal (Note scour pad at base of wall and 

increased stability berm) 

Reach NOE18 (Citrus Back Levee DM) 

This reach of levee is located between the floodwalls surrounding the Michoud Canal and 

Michoud Slip. There were no major renovations made to this reach or transition points. 

Reach NOE19 (Citrus Back Levee DM) 

This reach of floodwall is located around the Michoud Slip. The free standing height was 

reduced to approximately 6.0 feet (post-Katrina) from an average of 8.1 feet (pre-Katrina). A 

scour protection pad similar to those found in other figures was also placed at the base of the 

wall on the protected side. 

Reach NOE20 (Citrus Back Levee DM) 

This reach of levee is located between the floodwalls around the Michoud Slip and Bulk 

Loading Facility. There was some moderate scour damage to the levee, as shown in Figure NOE 

15, but no breaching. The only repairs made to this reach were for scour repair, as shown in 

Figure NOE 33. For the purposes of this risk assessment, it was assumed that the pre-Katrina and 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-39 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

post-Katrina conditions were the same. There were no scour protection improvements made to 

the levee/wall interfaces at the ends of this reach.

Figure NOE 33. Scour Damage Repairs Along Citrus Back Levee (Note: reference Figure NOE 15 which 

shows damage following Katrina) 

Reach NOE21 (Citrus Back Levee DM) 

This reach of floodwall failed during Katrina as shown in Figure NOE 16. It was replaced by 

an “L-wall,” which is assumed to have the same reliability as a “T-wall” within the context of 

this risk analysis. This new L-wall is shown in Figure NOE 34. Prior to Katrina, the previous Iwall had a weighted average top elevation of 12.1 (NAVD88 2004.65) and it is believed the 

failure was due to massive overtopping. The new L-wall was constructed to a top elevation of 

14.5 (NAVD88 2004.65).
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-40 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 35. New L-Wall at Bulk Loading Facility 

Reach NOE22 (Citrus Back Levee DM) 

This reach of levee is between the floodwalls that are east of the Amid Pump Station and the 

Bulk Loading Facility. There was minimal damage to this reach during Katrina and no major 

renovations were made by TFG, including no scour protection at the levee/floodwall interfaces. 

Reach NOE23 (Citrus Back Levee DM) 

This reach of I-wall is located just east of the Amid Pump Station. It was overtopped during 

Katrina and received minor scour damage. The free standing height of the wall was reduced to 

4.5 feet (post-Katrina) from 6.0 feet (pre-Katrina). A scour protection pad was also added at the 

base of the back side of the wall similar to those already shown in other areas. 

Reach NOE24 (Citrus Back Levee DM) 

This last reach of levee along the GIWW was damaged in a few spots during Katrina but had 

no major breaches. Major repairs undertaken by TFG within this reach included scour repairs 

around the Amid Pump Station discharge pipes along with scour protection improvements at the 

transition of levee/wall interfaces, including the heavily damaged area around the railroad 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-41 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

crossing closure near the GIWW/IHNC intersection. This area was heavily damaged during 

Katrina, as shown in Figure NOE 17. The scour protection repairs to this area are shown in 

Figure NOE 36, which illustrates a new transition wall on the east side of the closure and grouted 

rock at the end of the west transition wall to this closure.

Figure NOE 36. TFG Repairs to RR Closure at the GIWW/IHNC Intersection (Note: repairs include new 

transition wall and grouted rock) 

Reach NOE25 (IHNC East) 

This reach of levee was overtopped and damaged primarily at closure gate floodwall/levee 

interface transition points as shown in Figures NOE 18 and 19. Improvements made by TFG to 

this reach include providing grouted rock scour protection at levee/floodwall interfaces and 

repair of the major washout section that was damaged during Katrina. A typical scour 

improvement around a levee/floodwall interface along the IHNC is shown in Figure NOE 37. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-42 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Figure NOE 37. Transition Improvements along IHNC Levee/FW Interfaces 

Reach NOE26 (IHNC East) 

There were no major repairs made to this section of floodwall since the transition points are 

assigned to the adjoining levee reaches. 

Reach NOE27 (IHNC East) 

This relatively short reach of levee located between the I-10 and Highway 90 bridges was 

overtopped during Katrina with minimal damage. The only significant improvement to this reach 

was the placement of grouted rock at the levee/floodwall interfaces. 

Reach NOE28 (IHNC East) 

This section of floodwall is located behind the Folger’s Coffee Plant and continues to where 

it serves as the Dupuy Storage Building foundation. Scour protection, in the way of grouted rock, 

was provided at the base of the wall on the protected side. There is no significant change to the 

free standing height of the wall along this section as it was roughly 5.0 feet prior to Katrina. It is 

noted that the lower, unprotected old Highway 90 crossing was not significantly changed by 

TFG. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-43 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Reach NOE29 (IHNC East) 

This relatively short section of floodwall was placed as a separate reach since it serves as the 

building foundation (Dupuy Storage Facility), as shown in Figure NOE 20. It was made a 

separate reach since the building itself provided significant scour protection for this reach, but 

could also cause increased load without allowing overtopping. It was not damaged during 

Katrina and no major changes were made to this reach. 

Reach NOE30 (IHNC East) 

This last reach of the New Orleans East Basin consists primarily of concrete I-wall that was 

overtopped during Katrina. Several areas of the wall suffered significant scour on the protected 

side, see Figure NOE 21 for reference, and have been repaired by placing a scour pad at the base 

of the back side in the form of grouted rock. The free standing height of the I-walls along this 

reach has not been significantly changed from pre-Katrina condition since they averaged roughly 

6 feet at the time of the storm. 

Summary of Improvements - New Orleans East Basin 

There are a variety of improvements to the NOE Basin made by TFG. These primarily fall 

into two major categories: new walls/levees and scour protection. Scour protection 

improvements are primarily limited to the back side of floodwalls and transitions around 

levee/wall interfaces. However, where new levees were built along the southern portion of the 

East Levee and along the GIWW, improved materials and construction methods were used such 

that improved resistance to erosion from overtopping can also be expected when compared to the 

pre-Katrina condition. 

For the new floodwall sections, generally more reliable, stable walls replaced less reliable 

I-wall sections. This includes both T-wall and L-wall sections, which in the context of this 

analysis, are considered to be very reliable in terms of stability. In addition, these were built to 

an elevation considerably higher than the wall sections that failed, providing a higher level of 

protection for that reach. 

In an effort to make a relatively easy comparison of the physical changes to an individual 

NOE Basin reach from pre-Katrina to post-Katrina, Table NOE 1 was created. A general 

description of the reach location and number is provided along the left hand side of the table. For 

each reach, the type of protection, weighted average top elevation, wall free standing height 

(stick-up), and scour protection measures are shown in red for pre-Katrina conditions. The 

changes implemented by TFG to the reach characteristics are then shown in blue on the right 

hand side of the table. If there was no significant change to the reach for a particular 

characteristic(s), then the post-Katrina box is shaded gray. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-2-44 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers. 

Table NOE 1. Summary of Reach Changes for NOE Basin 

Reach Reach Description 

Pre-K 

Protect 

Pre-K 

Elev. 

Pre-K 

Stick-up 

Pre-K 

Scour 

Protect 

Post K 

Protect 

Post K 

Elev. 

Post-K 

Stick-up 

Post K Scour 

Protection 

NOE1 LF Airport I-wall 11.6 8.2 None Conc. Pad 

NOE2 Citrus LF Levee 13.3 n/a None

NOE3 Stars/Stripes West 

FW 

I-wall 13.5 10.0 None Conc. Pad 

NOE4 Stars/Stripes Levee Levee 13.3 n/a None

NOE5 Stars/Stripes East 

FW 

I-wall 13.7 7.5 None 6.0 Conc. Pad 

NOE6 Citrus LF Levee 12.9 n/a None

NOE7 Lincoln Beach 

Flood Wall 

I-wall 12.1 10.0 None 5.5 Conc. Pad 

NOE8 Citrus LF Levee 12.6 n/a None

NOE9 East Lakefront Levee 18.6 n/a None

NOE10 East Levee Levee 15.1 n/a None

NOE11 East Levee Levee 16.7 n/a None

NOE12 East Levee Levee 17.7 n/a None

NOE13 East Back Levee 15.5 n/a None New 

Levee 

19.5 Better Soils 

NOE14 PS #15 FW I-wall 17.5 7.5 None T-wall 22.1 n/a n/a 

NOE15 East Back Levee 16.8 n/a None New 

Levee 

19.5 Better Soils 

NOE16 Mich. Canal 

East FW 

I-wall 17.9 7.5 None 4.5 Conc. Pad 

NOE17 Mich. Canal West 

FW 

I-wall 20.8 9.3 None 7.0 Conc. Pad 

NOE18 Between Mich 

Canal and Slip 

Levee 17.2 n/a None

NOE19 Michoud Slip I-wall 16.7 8.1 None 6.0 Conc. Pad 

NOE20 Between Slip and 

Bulk Facility 

Levee 14.0 n/a None

NOE21 Bulk Load FW I-wall 12.1 6.0 None L-wall 14.5 n/a n/a 

NOE22 Levee Just West of 

Bulk Load FW 

Levee 13.4 n/a None

NOE23 FW East of Amid I-wall 14.5 6.0 None 4.5 Conc. Pad 

NOE24 Amid to IHNC Levee 13.8 n/a None

NOE25 GIWW to I-10 Levee 12.2 n/a None

NOE26 Under I-10 I-wall 12.4 6.0 None 5.0 Grout Rock 

NOE27 I-10 to Hwy 90 Levee 12.6 n/a None

NOE28 Hwy 90 to Dupuy 

Bldg 

I-wall 12.9 5.0 None Grouted Rock 

NOE29 Dupuy Bldg I-wall 13.5 6.0 Yes 

NOE30 Dupuy Bldg to 

Lakefront FW 

I-wall 12.2 5.5 None 5.0 Grouted Rock 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-1

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 3 

Jefferson Basin 

Field Reconnaissance and Definition of Reaches in Jefferson 

East Bank Basin 

The basin for Jefferson East Bank has been broken down into five distinct sections to 

develop both reaches and features definitions for the risk model. These sections are based on 

General Design Memorandum (GDMs) published by the New Orleans District, USACE and 

updated by field reconnaissance by IPET Team 10. The Jefferson East Bank Basin has been 

defined by the following sections shown in the figure below: 

1. West Return Levee 

2. Lakefront Levee 

3. 17th Street Canal 

4. Mississippi River Levees 

5. Jefferson/St. Charles Boundary Levees 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-2

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Overview of Jefferson Parish Basins 

Elevations 

All vertical elevations in this report are defined in NAVD88 2004.65 datum unless otherwise 

noted within the text. All horizontal datums are defined in State Plane Coordinates NAD83 – 

1702 Louisiana South, UTM NAD83 Zone 15, and GCS NAD83. 

Section 1: Kenner (West) Return Levee 

References 

The descriptions of the levees/floodwalls, stationing around the circumference of the polder, 

and terminology for associated features are referenced to the following document: 

DEPARTMENT OF THE ARMY, NEW ORLEANS DISTRICT, CORPS OF 

ENGINEERS, “Jefferson / St. Charles Parish Return Levee”, Design 

Memorandum No. 17a – General Design, New Orleans, LA, July 1987. 

Narrative 

The Kenner (West) Return Levee has a T-wall structure for the entire length along the St. 

Charles Parish border. This section begins at the northerly end of the runway for the New 

Orleans International Airport and ends at the Lake Ponchatrain levee. This section has been 

broken down into four subsections of incrementally increasing wall height. Reach 1 is from 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-3

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

station 0+00 W/L to 29+50 W/L is a T-wall section with an design elevation of 12.8. The 

elevation changes from station 29+50 W/L to 31+75 W/L, where the T-Wall crosses under I-10 

and the design elevation changes to 12.3 feet. The elevation of the T-wall changes back to design 

elevation 12.8 feet from station 31+75 W/L to 65+20.40 W/L. Reach 2 has a transition zone in 

the wall from station 65+20.40 W/L to 66+02.90 W/L where the T-wall design elevation 

increases to 13.3 feet. Reach 2 remains at the same elevation until station 130+70.00 W/L. At 

station 130+70.00 W/L, Reach 4 starts with another transition in the wall up to design elevation 

13.8 feet. The T-wall remains at elevation 13.8 feet until station 173+04.70 W/L. At station 

173+04.70 W/L, Reach 4 has a transition to design elevation 19.3 feet and continues until it 

connects with the Jefferson Lakefront Levee. 

The major damage in this section from Hurricane Katrina was scour at the base of the 

floodwalls and flood side slope protection, displacement of stone protection and slight movement 

and rotation of the T-walls. There were no failure in this section during the hurricane. The minor 

repairs to this section should be completed prior to 1 June 2006 however the repairs are not 

under the direction of Task Force Guardian. 

Typical T-Wall Transition in the Floodwall in Kenner Return Levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-4

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 1 – West return Levee – Summary of Reaches and Features 

Section Start End Structure EL 

Section / 

Point 

Length 

(feet) 

GDM Average 

Height (NVGD) 

1 00+00 0+55.0 T-Wall 

Transition 

El 13 – 

El 13.5 

S 55 13.25 

2 0+55.0 29+50.0 T-Wall El 13.5 S 2895 13.5 

3 29+50.0 31+75.0 Interstate 10 

Crossing 

El 11.5 P 225 11.5 

4 31+75.0 46+73.20 T-Wall El 13.5 S 1498 13.5 

5 46+73.20 47+09.20 T-Wall 

Transition 

El 13.5 – 

El 14.0 

P 36 13.75 

6 47+09.20 47+51.20 T-Wall at Pump 

Station 

El 14.0 P 42 14 

7 47+51.20 47+87.20 T-Wall 

Transition 

El 14.0 – 

El 13.5 

P 36 13.75 

8 47+87.20 65+20.40 T-Wall El 13.5 S 1733 13.5 Reach 1

9 65+20.40 66+02.90 T-Wall 

Transition 

El 13.5 – 

El 14.0 

P 83 13.75 

10 66+02.90 130+70.00 T-Wall El 14.0 S 6467 14 Reach 2

11 130+70.00 131+20.00 T-Wall 

Transition 

El 14.0 – 

El 14.5 

S 50 14.25 

12 131+20.00 173+04.70 T-Wall El 14.5 S 4235 14.5 Reach 3

13 173+04.70 178+54.70 T-Wall 

Transition 

El 14.5 – 

El 20 

P 549 17.25 

14 178+54.70 180+69.70 T-Wall El 20 S 215 20 

15 180+69.70 180+74.70 T-Wall El 20.5 S 5 20.5 Reach 4

Points Width Elevation 

1 I-10 Crossing 147 11.5 

2 Parish Line 

Pump Station 

114 14 

3 Gate W-7 West 

Esplanade 

6 14 

4 Gate W8 Vintage 

Street 

6 14 

Section 1 – West Return Levee - Summary Coordinates for Defined Reaches 

State Plane NAD83 UTM NAD 83 

1702 - LA South 

Northing Easting Graphical Latitude Longitude 

Zone 15 

Northing Easting 

Return Levee

Reach 1 START 546407.659 3614397.566 29 59 53.87687 90 16 46.29803 2501430.19 10897988.09 

Reach 1 END 

Reach 2 START 

552907.875 3614407.909 30 00 58.22085 90 16 45.50041 2501345.82 10904491.17 

Reach 2 END 

Reach 3 START 

558810.683 3614388.995 30 01 56.65366 90 16 45.09787 2501240.83 10910396.17 

Reach 3 END 

Reach 4 START 

563727.833 3614373.814 30 02 45.32911 90 16 44.75590 2501153.88 10915315.15 

Reach 4 END 564277.527 3614378.32 30 02 50.77003 90 16 44.64708 2501150.37 10915865.14 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-5

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Definition of reaches 

Reach 1 – This reach consists of T-wall at average elevation 10.9 feet with a length of 

approximately 6,520 feet. There are two features in this reach. The features are the I-10 

floodwall at elevation 10.4 feet and the Parish Line Pumping Station at elevation 11.75 feet. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-6

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

I-10 Floodwall 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-7

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Parish Line Pump Station (Note: Pump discharge through T-wall) 

Reach 2 – This reach consists of T-wall at average elevation 11.0 feet with a length of 

approximately 6,550 feet. The only feature in this reach is Gate W7, West Esplanade Ave. The 

swing gate is 6 feet wide and 8 feet high with a sill elevation of 4.3 feet. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Gate W-7 West Esplanade 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-9

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 3 – This reach consists of T-wall at average elevation 12.6 feet with a length of 

approximately 4,285 feet. There is only one feature in this reach is Gate W8, Vintage Ave. The 

swing gate is 6 feet wide and 8 feet high with a sill elevation of 4.3 feet. 

Gate W8 Vintage Street 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-10

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical T-wall and slope protection on floodside in Reach 3 

Reach 4 – This reach consists of T-wall at average elevation 15.7 feet with a length of 

approximately 769 feet. There are no features in this reach but there is a sheet pile transition into 

the Lakefront levee. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-11

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 2: Jefferson Parish Lakefront Levee 

References 

Descriptions of the levees/floodwalls, stationing around the circumference of the polder, and 

terminology for associated features are referenced to the following documents: 

DEPARTMENT OF THE ARMY, NEW ORLEANS DISTRICT, CORPS OF 

ENGINEERS, “Jefferson Parish Lakefront Levee”, Design Memorandum No. 17 

– General Design, Volumes I and II, New Orleans, LA, November 1987. 

Narrative 

The Jefferson Parish Lakefront Levee consists primarily of levee sections intermingled with 

four pump stations, two boat launch facilities and the Pontchartrain Causeway. This section has 

been broken down into four subsection (Reaches 5 to 8) defined on changes in levee elevation 

and the location of the pumping stations. Reach 5 is an earthen levee from station 0+00 B/L to 

128+00.00 B/L with a crest elevation originally constructed at 17.3 feet, a 10-foot crest width, 

and 1 on 3 vegetated side slopes. Pump Station 4 (Duncan Canal PS) is located between stations 

115+00.00 B/L and station 128+00.00 B/L. The original pump station has been taken out of 

service and replaced by a new pump station that is situated just westerly of the old station. 

T-walls have been reconfigured around the old and new pump stations and transition into the 

levee sections on both sides. The design elevation of the original Pump Station 4 is at elevation 

21.7 feet. It appears that the walls surrounding the new pump station are at the same elevation of 

the old pump station. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-12

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

There was no visible damage from Hurricane Katrina in this reach. Most damage was to the 

foreshore protection but little or no damage occurred to the levee or at the pump station walls. 

There will be only minor repairs that need to be completed prior to 1 June 2006. 

Foreground: Original Pump Station and old masonary wall (not in Service) 

Background: Duncan Pump Station with new T-walls 

Reach 6 runs from Pumping Station No. 4 at station 128+00 B/L to Pumping Station No. 3 at 

station 210+00 B/L. The earthen levee cross section is similar to that in Reach 5 from station 

128+00 to 141+00 B/L. From station 141+50 B/L to 145+50 B/L, the crest changes to design 

elevation of 15.8 to provide access to the Williams Boulevard Boat Launch. The elevation of the 

ramp over the levee is at elevation 13.8. From station 146+00 B/L to 184+50 B/L, the levee cross 

section is again typical to that on the other side of the boat launch. From station 185+50 B/L to 

Pumping Station No. 3 (Elmwood Canal) at station 210+00 B/L, the levee crest was originally 

designed to elevation 15.3 feet. Cantilever sheeting forms the transition from the levee to both 

sides of the pumping station. Along the lakefront, damage to the foreshore concrete slab 

protection was evident. It is noted that Pumping Station No. 3 consists of original construction 

on the western side and a later expansion on the eastern side that appears to have been built in 

the 2001 time frame. 

There was no visible damage from Hurricane Katrina in this reach. Most damage was to the 

foreshore protection but little or no damage occurred to the levee or at the pump station walls. 

There will be only minor repairs that need to be completed prior to 1 June 2006. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-13

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 7 runs from Pumping Station No.3 at station 221+45 to Pumping Station No.1 at 

station 464+50 B/L. The levee cross section is typical throughout the reach with a crest elevation 

originally constructed to elevation 15.3 feet and vegetated side slopes. Cantilever sheeting forms 

the transition from the levee to Pumping Station No. 2. It is noted that Pumping Station No. 2 

consists of original construction on the western side and a later expansion on the eastern side that 

appears to have been built in the 2001 time frame. On the west side of this pumping station, the 

crest at the access road is approximately two feet low to facilitate passage of vehicles and the top 

elevation of the transition sheeting is lower than the crest of the levee. A portion of this reach 

was being re-constructed at the time of this reconnaissance inspection. Reference contract 

number W912P8-05-C-0014 for specific details. At the time of the inspection, there was a 

200-foot gap in the levee near Pumping Station No. 2 that the contractor was using for access. 

This gap needs to be filled and vegetation needs to be established in order to restore the integrity 

of this reach. It is noted that the Lake Pontchartrain Causeway Boulevard intersects Reach 7 

between station 434+80 B/L to 438+40 B/L. Existing details at the causeway are shown on the 

original design memorandum, which indicates two bottom roller gates to close the line of 

protection with the causeway below the crest of the levee. This section has changed and the 

causeway now passes over the protection structures, and the floodwalls form the transition from 

the causeway to the levee. 

There was no visible damage from Hurricane Katrina in this reach. Most damage was to the 

foreshore protection but little or no damage occurred to the levee or at the pump station or 

Causeway walls. There will be only minor repairs that need to be completed prior to 

1 June 2006. 

Reconstruction of a Portion of Reach 7 Lakefront Levee 

Reach 8 runs from Pumping Station No. 1 at station 479+95 B/L to its junction with the 

17th Street Outfall Canal at station 550+22 B/L. From station 479+95 B/L to 485+00 B/L, the 

levee was originally constructed to elevation 15.7 feet. The access road to the Bonnabel Boat 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-14

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Launch Area intersects the levee between stations 485+00 B/L to 486+50 B/L. The net grade for 

the access road is elevation 13.8 feet, and two swing gates are installed to maintain the level of 

protection. The top of the swing gates appears to be six inches lower than the top of the levee. 

After the access road to the boat launch, the levee transition from elevation 15.7 feet at station 

486+50 to elevation 15.3 feet at station 487+50 and continues at this elevation to the limit of 

work at station 550+22. 

There was no visible damage from Hurricane Katrina in this reach. Most damage was to the 

foreshore protection but little or no damage occurred to the levee or at the pump station walls. 

There will be only minor repairs that need to be completed prior to 1 June 2006. 

Section 2 – Jefferson lakefront levee - Summary of Reaches and Features 

Section Start End Structure EL 

Section/ 

Point Length

GDM Average 

Height 

1 0+85 1+45 Levee Transition El 13.0 - El 

16.0 

S 60 15.5 

2 1+45 115+00 Levee El 16 S 11005 16 

3 115+00 128+00 I-Wall/T-Wall for Pump 

Station #4 

El 17.0 - El 

22.5 

P 1300 18.25 Reach 5

4 128+00 141+00 Levee El 16 S 1300 16

5 141+00 141+50 Levee Transition El 16 - El 

14.5 

P 50 17.25

6 141+50 145+50 Levee - William Blvd Boat 

Launch 

El 14.5 P 400 14.5

7 145+50 146+00 Levee Transition El 14.5 - El 

16.0 

P 50 15.25

8 146+00 208+00 Levee El 16 S 6200 16 Reach 6

9 208+00 223+00 I-Wall for Pump Station #3 El 16.5 P 1500 16.5

10 223+00 343+00 Levee El 16 S 12000 16

11 343+00 355+00 I-Wall for Pump Station #2 EL 16.5 P 1200 16.5

12 355+00 434+80 Levee EL 16 S 7980 16

13 434+80 438+40 Causeway Blvd El 16.5 P 360 16.5

14 438+40 464+50 Levee El 16 S 2610 16 Reach 7

15 464+50 479+95 I-Wall/T-Wall for Pump 

Station #1 

El 16.5 - El 

22.5 

P 1540 19.5

16 479+95 487+50 Levee - Bonnabel Boat 

Launch 

El 14.5 P 760 14.5

17 487+50 550+22 Levee El 16 S 6270 16 Reach 8

Points Width Elevation 

1 1+00 Gate L1 Floodwall 20 17 

2 104+22 Gate L3 Duncan Canal 22 17 

3 145+13 Gate L4 Williams Blvd 60 14.5 

4 402+04 Gate 2 - Causeway SB (no longer used) 45 14.5 

5 402+45 Gate 3 - Causeway NB (no longer used) 45 14.5 

6 475+54 Gate L9A Bonnabel Blvd South 22 16 

7 475+54 Gate L9B Bonnabel Blvd North 22 16 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-15

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 2 – Jefferson Lakefront Levee - Summary Coordinates for Defined Reaches 

State Plane NAD83 UTM NAD 83 

1702 - LA South 

Northing Easting 

Graphical 

Latitude Longitude 

Zone 15 

Northing Easting 

Lakefront Levee

Reach 5 START 564277.527 3614378.32 30 02 50.77003 90 16 44.64708 2501150.37 10915865.14 

Reach 5 END 

Reach 6 START 

561991.187 3625823.549 30 02 27.07773 90 14 34.66901 2512633.87 10913744.89 

Reach 6 END 

Reach 7 START 

559141.211 3633246.774 30 01 58.15966 90 13 10.52729 2520102.00 10911001.96 

Reach 7 END 

Reach 8 START 

554468.982 3656812.867 30 01 09.56725 90 08 42.98123 2543747.54 10906671.25 

Reach 8 END 555079.333 3664232.118 30 01 14.83996 90 07 18.51511 2551161.68 10907390.11 

Definition of Reaches 

Reach 5 – This reach consists of a levee at average elevation 15.8 feet with a length of 

approximately 12,365 feet. There are two features in this reach. The first is Gate L1 at the start 

of the Lakefront Levee near the West Return Levee. Gate L1 is 20 feet in width and has a sill 

elevation of 10.0 feet. The second feature is the Duncan Canal Pumping Station at wall elevation 

20.0 feet. There is a T-wall section around the new pump station which is approximately 

2,236 feet in length. There is a swing gate at station 104+22 near the pump station, Gate L3, 

which is 22 feet wide and has a sill elevation at 10.0 feet. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-16

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Gate L1 – Lakefront Levee 

Gate L3 – Duncan Canal Pump Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-17

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Duncan Canal Pump Station 

Reach 6 – This reach consists of Levee at average elevation 14.1 feet with a length of 

approximately 8,000 feet. There one feature in this reach which is the William Blvd Boat Launch 

at with a ramp at elevation 13.8 feet. The ramp also has a gate at station 145+13 which is 60 feet 

in width and a sill elevation at 9.25 feet. 

Reach 7 – This reach consists of a levee at an average elevation 14.3 feet and length of 

approximately 25,650 feet. There two features contained in this reach. The first feature is Pump 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-18

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Station #3 - Elmwood Canal at elevation 18.0 feet. The pump station has an uncapped sheet pile 

wall of about 750 feet in total length. The second feature in this reach is Pump Station #2 - 

Suburban Canal at elevation 18.0 feet. This pump station has an uncapped sheet pile wall of 

about 864 feet in total length. The Pontchartrain Causeway Boulevard is not considered a feature 

since the lanes have been raised above the protection walls and levees. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-19

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pump Station #3 – Elmwood Canal Pump Station 

Pump Station #2 – Suburban Canal Pump Station 

Reach 8 – This reach consists of levee at average elevation 14.4 feet with a length of 

approximately 8,570 feet. There two features contained in this reach. The first is the Pump 

Station #1- Bonnabel Canal at elevation of 18.0 feet. This pump station has a sheet pile transition 

wall into a I-wall/T-wall of about 1,052 feet in total length. The second feature is Bonnabel Boat 

Launch which has a elevation 15.3 feet. The boat launch has two gates, L9A and L9B, which 

have a width of 22 feet and a sill elevation of 11.8 feet. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-20

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pump Station #1 – Bonnabel Canal Pump Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Bonnabel Boat Launch Ramp and Gate 

Section 3: 17th Street Outfall Canal 

References 

Descriptions of the levees/floodwalls, stationing around the circumference of the polder, and 

terminology for associated features are referenced to the following documents: 

DEPARTMENT OF THE ARMY, NEW ORLEANS DISTRICT, CORPS OF 

ENGINEERS, “Orleans Parish, Jefferson Parish, 17th Street Outfall Canal 

(Metairie Relief) Design Memorandum No. 20 – General Design, Volumes I and 

II, New Orleans, LA, March 1990. 

Narrative 

The west line of the 17th Street Outfall Canal runs from station 0+00 at its junction with the 

Lakefront Levee to station 119+95 at Pumping Station No. 6. This section is constructed 

primarily of I-wall sections on top of earthen levees. The west line of the 17th Street Canal has 

been broken out into three reaches based on elevation changes along the canal. Reach 9 has a 

design elevation 13.3 feet from stations 0+00 to 75+70 at the Veterans Highway Bridge. Reach 

10 has a design elevation 13.8 feet from stations 77+70 to 92+50. Reach 11 has a design 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-22

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

elevation of 14.3 feet from stations 93+50 to 119+95. At the junction of the Lakefront Levee and 

the west wall of the 17th Street Outfall Canal north of Hammond Highway, construction work 

was ongoing under contract number W912P8-06-C-0008, titled “17th Street Canal Interim 

Closure Structure.” Interstate Highway 10, which passes over the I-wall in the vicinity of 

stations 90+00 to 92+50, has no effect on the integrity of the hurricane protection. There is a gate 

structure in the west wall just to the north of Pumping Station No.6 through which the CSX 

railroad line passes. The area was littered with sandbags indicating that the swing gate may not 

seal tightly. No storm related damage was observed along the west wall. 

Section 3 – 17th Street Outfall Canal - Summary of Reaches and Features 

Section Start End Structure EL 

Section / 

Point Length

GDM Average 

Height 

1 0+00 0+10 Floodwall 16.5 P 10 16.5 

2 0+10 0+43.60 Floodwall 

Transition 

El 16.5 – 

El 14.0 

S 34 15.25 

3 0+43.60 2+75.10 I-Wall El 14.0 S 231 14 

4 2+75.10 2+81.0 T-Wall El 14.0 P 6 14 

5 2+81.0 3+05.10 Gate No. 1 El 14.0 P 24 14 

6 3+05.10 3+11.0 T-Wall El 14.0 P 6 14 

7 3+11.0 3+40.50 I-Wall El 14.0 S 29 14 

8 3+40.50 3+47.00 T-Wall El 14.0 P 7 14 

9 3+47.00 4+09.00 Gate No. 2 El 14.0 P 62 14 

10 4+09.00 4+15.50 T-Wall El 14.0 P 7 14 

11 4+15.50 75+70.00 I-Wall El 14.0 S 7154 14 Reach 9 

12 75+70.00 76+00.00 I-Wall Transition El 14.0 – 

El 14.5 

S 30 14.25 

13 76+00.00 76+10.04 I-Wall El 14.5 P 10 14.5 

14 76+10.04 77+22.04 Veterans Hwy El 14.5 S 112 14.5 

15 77+22.04 92+50.0 I-Wall El 14.5 S 1528 14.5 Reach 10

16 92+50.0 93+50.0 I-Wall Transition El 14.5 – 

El 15.0 

S 100 14.75 

17 93+50.0 119+95.49 I-Wall El 15.0 S 2645 15 Reach 11

Points Width Elevation 

1 Gate L 10 Orpheum Ave - 58 ft gap (under construction) 58 9.5 

2 Old Hammond Hwy Bridge - 47 ft gap (under construction) 47 10 

3 Gate E4 - Veterans Blvd. 8 14.5 

4 Gate E5 - Veterans Blvd 8 14.5 

5 Gate E8 - Canal Street 10 15 

6 Gate E9 - Southern RR 22 15 

7 Pump Station #6 150 14.5 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-23

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 3 – 17th Street Outfall Canal - Summary of Coordinates for Defined Reaches 

State Plane NAD83 UTM NAD 83 

1702 - LA South 

Northing Easting 

Graphical 

Latitude Longitude 

Zone 15 

Northing Easting 

17th St. Canal Levee

Reach 9 START 555079.333 3664232.118 30 01 14.83996 90 07 18.51511 2551161.68 10907390.11 

Reach 9 END 

Reach 10 START 

547700.195 3663758.46 30 00 01.84526 90 07 24.78918 2550795.34 10900000.24 

Reach 10 END 

Reach 11 START 

545946.804 3663664.746 29 59 44.49864 90 07 26.06543 2550727.12 10898244.58 

Reach 11 END 542850.391 3663595.224 29 59 13.85519 90 07 27.22762 2550702.65 10895145.55 

Definition of Reaches 

Reach 9 – This reach consists of an I-wall in levee section at an average elevation of 

12.5 feet with a length of approximately 7,570 feet. The reach actually starts at the Lakefront 

levee however this area was under construction for flood proofing of the Hammond Highway 

bridge and was not accessible for inspection. There are two main features in this reach, Gate 1 at 

elevation 14.0 feet and Gate 2 is at elevation 14.0 feet. However, these gates have been removed 

for the construction work at the Hammond Highway. The sequence of pictures below defines the 

linear west line of Reach 9 (Note: These pictures are pre-Katrina). 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-24

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-25

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 10 – This reach consists of an I-wall in a levee at an average elevation of 12.9 feet 

with a length of approximately 1,680 feet. This reach goes from Veterans Boulevard to just south 

of the I-10 overpass. There are two features in this reach, Gate E4 and Gate E5 with a elevation 

of 12.9 feet and a sill elevation of 6.5 feet. The Veterans Boulevard and I-10 overpass bridges 

have been floodproofed with tie-in parapet walls higher than the canal I-walls and are not 

considered features. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 11 – This reach consists of I-wall in a levee at an average elevation of 13.4 feet with a 

length of approximately 2,745 feet. This reach goes from the I-10 overpass to Pump Station #6. 

There three features in this reach. The first is Canal Street Gate E8 which is at elevation of 

12.0 feet. The second feature is the Canal Street Pump Station at elevation 13.5 feet. The final 

feature is the Southern Railroad Swing Gate E9 at elevation 10.3 feet. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-30

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-31

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-32

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pump Station #6 

Gate E9 CSX Railroad 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-33

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 4: Mississippi River Levees 

References 

General design information was not available for the MRT levees within Jefferson Parish at 

the time of the report. Information was gathered from discussion with the East Jefferson Parish 

Levee Board and from the field recon. 

Narrative 

The MRT levees are earth structures with a 10-foot crown and side slopes of varying grades. 

About half of the levees along this section of the river have concrete scour protection on the 

flood side. There is no scour protection on the protected side of the levees. The levees are also 

well maintained with both an asphalt road and mowed grass slopes on both sides. The crown of 

the levee is extensively used by the public as a bike and walking path. There are a number of 

pipeline crossings and access ramps that were defined for his section. However none of these 

structure greatly interfered with the integrity of the levees so they were not include into the risk 

assessment. 

This section as been broken down into three reaches defined on the change in elevation of the 

levees. Reach 12 starts at the Orleans East Parish line and goes along the river from station 0+00 

to 388+17.51. Most of this reach was unprotected by the concrete scour protection. Reach 13 

starts at station 388+17.51 and continues to station 538+17.51. This reach has concrete scour 

protection on about two-thirds of the flood side banks. Reach 14 starts at station 538+17.51 and 

runs to the St. Charles Parish border just southwest of the international airport. Most of this 

reach has concrete scour protection on the flood side banks. 

Flood side of MRT in East Jefferson Parish 

(Note: Concrete scour protection and asphalt roadway on crown) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 4 – Mississippi River Levees - Summary of Reaches and Features 

MRT Levees 

1 0+00 388+17.51 Levee 26.6 S 38818 26.6 Reach 12 

2 388+17.51 538+17.51 Levee 25 S 15000 25 Reach 13 

3 538+17.51 608+12.06 Levee 24 S 6995 24 Reach 14 

Section 4 – Mississippi River Levees - Summary of Coordinates for Defined Reaches 

State Plane NAD83 UTM NAD 83 

1702 - LA South 

Northing Easting 

Graphical 

Latitude Longitude 

Zone 15 

Northing Easting 

Mississippi Levee 

Reach 12 START 530021.436 3659583.755 29 57 07.28044 90 08 14.36542 2546875.73 10882251.64 

Reach 12 END 

Reach 13 START 

532634.821 3653607.193 29 57 33.76261 90 09 21.99695 2540858.17 10884779.48 

Reach 13 END 

Reach 14 START 

523758.1 3642759.654 29 56 06.97997 90 11 26.31882 2530134.31 10875740.84 

Reach 14 END 537504.405 3614269.527 29 58 25.75591 90 16 48.68478 2501431.68 10889079.25 

Reach 12 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-35

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 13 

Reach 14 

Section 5: Jefferson/St. Charles Parish Border Interior Levees 

References 

General design information was not available for the MRT levees within Jefferson Parish at 

the time of the report. Information was gathered from discussion with the East Jefferson Parish 

Levee Board and from the field recon. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-36

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Narrative 

The MRT levees are earth structures with a 10-foot crown and side slopes of varying grades. 

About half of the levees along this section of the river have concrete scour protection on the 

flood side. There is no scour protection on the protected side of the levees. The levees are also 

well maintained with both an asphalt road and mowed grass slopes on both sides. The crown of 

the levee is extensively used by the public as a bike and walking path. There are a number of 

pipeline crossings and access ramps that were defined for his section. However none of these 

structure greatly interfered with the integrity of the levees so they were not include into the risk 

assessment. 

This section as been broken down into three reaches defined on the change in elevation of the 

levees. Reach 12 starts at the Orleans East Parish line and goes along the river from station 0+00 

to 388+17.51. Most of this reach was unprotected by the concrete scour protection. Reach 13 

starts at station 388+17.51 and continues to station 538+17.51. This reach has concrete scour 

protection on about two-thirds of the flood side banks. Reach 14 starts at station 538+17.51 and 

runs to the St. Charles Parish border just southwest of the international airport. Most of this 

reach has concrete scour protection on the flood side banks. 

Section 5 – Jefferson/St. Charles Parish Border Interior Levee - Summary of Reaches 

and Features 

Levee from Airport to MRT Not Needed Since St Charles Has New Levee To Tie To Airport Runway Extension 

1 0 1000 Gap in levee 5 S 1000 5 

2 1000 1485 Sheet pile wall 10 S 485 10 

3 1485 1735 Gap in levee 3.5 S 250 3.5 Reach 15 

4 0+00 32+89 Levee 10 S 3289 10 

5 32+89 42+69.9 Sheet pile wall - 42 ft gap @ 5.9 ft at RR 10.8 S 980 10.8 Reach 16 

6 42+69.9 57+00 Sheet pile wall 10.5 S 1431 10.5 Reach 17 

7 57+00 74+00 Levee 13 S 1700 13 Reach 18 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-37

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 5 – Jefferson/St. Charles Parish Border Interior Levee - Summary of 

Coordinates for Defined Reaches 

State Plane NAD83 

1702 - LA South Graphical 

UTM NAD 83 

Zone 15 

Northing Easting Latitude Longitude Northing Easting 

Return Levee 

Reach 15 START 537504.405 3614269.527 29 58 25.75591 90 16 48.68478 2501431.68 10889079.25 

Reach 15 END 

Reach 16 START 

539488.473 3614278.54 29 58 45.39526 90 16 48.37495 2501411.83 10891064.28 

Reach 16 END 

Reach 17 START 

543565.735 3614781.939 29 59 25.70992 90 16 42.22427 2501856.10 10895150.57 

Reach 17 END 

Reach 18 START 

544766.352 3614391.643 29 59 37.63026 90 16 46.53705 2501448.16 10896346.01 

Reach 18 END 546407.659 3614397.566 29 59 53.87687 90 16 46.29803 2501430.19 10897988.09 

Reach 15 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-38

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 16 

Reach 17 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-39

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 18 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-40

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Jefferson West Bank Area 

The West Bank Basin is composed of four sub-basins that are designed as three projects. 

These are 1) Cataouatche, 2) Westwego to Harvey Canal, and 3) Harvey Canal to Algiers Canal. 

Cataouatche – JW1 and JW2 

This area is located in Jefferson Parish and is generally bounded by the Mississippi River and 

its alluvial ridge to the north and the Lake Cataouatche levee to the west, south and east. The 

topography is flat with ground elevations ranging from +7.5 feet NGVD on the alluvial ridges 

along the Mississippi River to –5 feet NGVD in the interior of the area. Approximately 

40 percent of the area is below sea level. The surface area is 22.6 square miles. The area is 

protected by 25.4 miles of levees, natural ridges and floodwalls. 

Segment 1 extends from the main line Mississippi River levee (MRL) at the Jefferson Parrish 

boundary southward to the Texas and Pacific railroad tracks. There are no levees or dikes in this 

area. The natural contour of the area provides the protection, but this segment is listed since it is 

possible for storm surges to flank the Segment 2 levee reach and cause flooding. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-41

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Segment 2 is the proposed levee that follows the crushed stone roadway southward from the 

Texas and Pacific railroad tracks (that becomes an asphalt roadway which is used by the land fill 

operators in the area) to US 90. 

Texas and Pacific railroad 

Local levee along site of proposed HPS levee 

Segment 3 is a short, small dike built parallel to Hwy 90. Hwy 90 is a 4-lane road with a 

raised median in the center. The median provides the higher level of protection. The road rises as 

a low relief ramp at the beginning of Segment 4. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-42

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Segment 4 is an earthen levee extending southward from Hwy 90 to the Cataouatche 

Pumping Station. The discharge lines of the first pumping station pass over the levee. The 

discharge lines of the second pump station (immediately adjacent to the first station) pass 

through a sheetpile wall. 

Levee begins at Hwy 90 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-43

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical levee in this area 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-44

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pump station near Hwy 90 with pipe crossing over levee 

Segment 5 is an all clay levee that extends eastward from the Cataouatche Pumping Station 

to the I-Wall in the Segnette State Park. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-45

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Cataouatche Pump Station 

Sheetpile wall transition to concrete capped I-wall at the Segnette State Park 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-46

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Segment 6 is a concrete I-Wall atop a clay levee. The controlling grade listed for this area is 

the preconstruction levee grade. The area has two vehicular gates. 

Concrete capped I-wall at the Segnette State Park 

Swing in the concrete capped I-wall at the Segnette State Park where wall ends at the Segnette Pump 

Station 

Segment 7 is completed floodwalls that lie between the Segnette Pump Station and the Old 

Westwego Pump Station. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-47

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Segnette Pump Station 

Segment 8 extends from the floodwall at the head of Company Canal (closest line of flood 

protection to the Mississippi River) to the MRL. The natural contour of the area provides the 

protection. 

Segment 9 is the West Jefferson Levee District Mississippi River levee. This all clay levee 

closes the north end of the sub-basin and extends from Westwego to the St. Charles Parish line. 

Segment 10 is the interior drainage separator, which begins at the end of Segment 3. It 

proceeds along US 90 to the east until it intersects the Texas and Pacific railroad tracks just north 

of the Westbank Expressway. It then continues to the east along the railroad until it intersects 

Segment 8 and then turns north to the MRL. JW1 is to the north of the Segment 10 interior levee 

and JW2 is to the south. 

There are a total of five vehicular floodgates (double swing) and two pedestrian (single 

swing) floodgates in the protection system. The sill elevations of these floodgates are at or above 

the current controlling elevation so these gates are not a factor in draining the area. 

There are four pumping stations that drain the protected area.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-48

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Westwego to Harvey Canal – JW3 

This area is located in Jefferson Parish and is generally bounded by the Mississippi River and 

its alluvial ridges on the north, the Harvey Canal on the east and marshes/wetlands on the south 

and west. The topography is flat with ground elevations ranging from +7.5 feet NGVD on the 

alluvial ridges along the Mississippi River to –4 feet NGVD in the interior of the area. 

Approximately 40 percent of the area is below sea level. The surface area is 21.4 square miles. 

The area is protected by 27.5 miles of levees and floodwalls. 

Segment 1 is a floodwall stretching between the Old and New Westwego Pumping Stations 

and connects the Cataouatche sub-basin to Westwego to Harvey Canal sub-basin. The 

segregation of these two sub-basins is not very pronounced. The general contour tie to the 

Mississippi River levee is described in Segment 8 the JW1 Cataouatche sub-basin. 

Segment 2 is the Westwego Levee that is a geosynthetic reinforced, clay levee running 

parallel to Mayronne Canal between the New Westwego Pumping Station and Dugues CanalWestwego Seaplane Airport. A 400-ft canal closure was constructed at the head of the Dugues 

Canal.

Segment 3 runs between Dugues Canal and the New Westminster Pump Station and the 

North-South Levee. This levee is all clay. 

Segment 4 is the Westminster Levee, which parallels the Grand Cross Canal, stretches 

between New Westminster Pumping Station and Orleans Village Pumping Station (out of 

service). This clay levee is geosynthetically reinforced. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-49

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

New Westminster Pump Station 

Segment 5 is the Orleans Village levee, which is all clay and paralleling Glasco Canal, 

between Orleans Village Pumping Station (out of service) and Oak Cove Pumping Station. 

Along this reach is the Ames and Mount Kennedy Pumping Stations connected by floodwall. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-50

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Ames Pump Station 

Segment 6 consists of the Oak Cove and Hwy 45 clay levees running between Oak Cove 

Pumping Station and the Hwy 45 crossing. Also found along this length are areas of T-wall, 

I-wall, and one vehicular floodgate at Hwy 45. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-51

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Sheet pile transition at the LA Hwy 45 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-52

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Double Swing gate at the LA Hwy 45 closure in the V-Line levee 

Segment 7 is the V-Line Levee which is an I-wall between LA Hwy 45 and Hwy 3134. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-53

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

V-line levee continues south of LA Hwy 45 closure 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-54

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Southern tip of V-line levee 

Segment 8 stretches from the V-Line Levee floodwall to the Old Estelle Pumping Station and 

is an all clay levee with one main road crossing. 

Segment 9 is an all clay levee running parallel along the North bank of the Old Estelle 

Pumping Station Outfall Canal. It runs to the Harvey Canal. 

Segment 10 is the West Bank Harvey Canal Levee. It consists of a clay levee running from 

the mouth of the Harvey Canal to the LaPalco bridge. Along this segment is the New Estelle 

Pumping Station, a floodwall at the Bridgeline pipeline, and three areas of sheetpile closure 

required because of unstable earthen levee sections. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-55

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Sheetpile closures 

New Estelle Pump Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-56

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Lapalo Bridge Overpass construction of sector gate under way 

Segment 11 stretches from LaPalco Bridge to the Harvey Lock, paralleling the Harvey Canal. 

This floodwall includes the Harvey and Cousins Pumping Stations, a vehicular gate and ties the 

Westwego and Harvey Canal sub-basin back into the Mississippi River Main Line levee. 

Industrial area along Harvey Canal 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-57

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

US 90 Bridge over Harvey Canal near Harvey Lock 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-58

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Looking south down Harvey Canal from Hwy 45 Bridge near Harvey Lock 

Harvey Lock 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-59

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Harvey Lock exit toward MS River 

MS River Levee at exit of Harvey Canal 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-60

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Harvey Pump Station and surrounding walls 

Segment 12 is the West Jefferson Levee District Mississippi River levee. It encloses the 

north side of the sub-basin between Westwego and Harvey Canal and is an all clay levee. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-61

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical MRL. Paved 10 foot crown. Armor on flood side 

MRL in vicinity of Northrup-Grumman Plant 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-62

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical closure gate (vehicle and pedestrian) along 

MRL 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-63

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

There are 10 pedestrian floodgates (swing) and two roadway floodgates (one swing and one 

miter). 

There are 11 pumping stations that drain the protected area. The location of the pumping 

stations.

Harvey Canal to Algiers Canal – JW4

This area is located in Jefferson and Plaquemines Parish and is generally bounded by the 

Mississippi River on the north, the Jefferson, Plaquemines & Orleans Parish lines on the east, the 

Algiers Canal on the south, and the Harvey Canal on the west. The topography is flat with 

ground elevations ranging from +15 feet NGVD on the alluvial ridges along the Mississippi 

River to -5 feet NGVD in the interior of the area. Approximately 40 percent of the area is below 

sea level. The surface area is 18.8 square miles. The area is protected by 21.3 miles of levees and 

floodwalls. 

Segment 1 extends from the Harvey Canal Lock at the Mississippi River down the East bank 

of the Harvey Canal to the Hero Pumping Station where the pumping station discharge lines pass 

through a T-wall. This clay levee is a local levee in a heavily industrialized area. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-3-64

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Segment 2 extends from the South end of the Hero Pumping Station around the bend where it 

ties into the Algiers Canal levee. The clay levee is also a local levee in a heavily industrialized 

area. 

Segment 3 picks up where segment 2 ended and continues along the West bank of the 

Algiers Canal. The clay levee is interrupted by floodwall segments that cross over the Belle 

Chasse tunnel and in front of Planters Pumping Station. It ends at the tie-in of the local levee 

separating Plaquemines and Orleans Parishes. A railroad track crosses over the top of the 

existing levee. A future floodgate is planned for the area. 

Segment 4 is an all clay levee that runs along the length of the Orleans Parish line between 

Algiers Canal and the Mississippi River levee at the Greater New Orleans Bridge. 

Segment 5 is the West Jefferson Levee District Mississippi River Levee stretching between 

the Harvey Canal and the Orleans Parish line beneath the Greater New Orleans Bridge. This 

levee consists of all clay levees with short reaches of concrete I-Wall atop clay levees with 

railroad and vehicular gates. 

There are no floodgates, control structures, or drainage structures in the protection system. 

There are two pumping stations that drain the protected area.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-1

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 4 

St. Charles Basin 

The St. Charles hurricane protection system (HPS), shown in Figure 1, was designed as part 

of the Lake Pontchartrain, LA, and Vicinity Hurricane Protection Project. The St. Charles HPS 

protects 17.2 square miles of urban, industrial, commercial, and ecological lands that is 

essentially a low density residential community with a small business district along U.S. 

Highway 61. The St. Charles Basin is generally bounded on the north by the St. Charles HPS, on 

the south by the Mississippi River Levee (MRL) and on the west by the Bonnet Carre guide 

levee. As designed, the HPS levees were generally constructed with a 10-foot crown width with 

side slopes of 1V on 3H for both the flood side and protected side. Topography is flat with 

ground elevations ranging from +12 feet NGVD on the alluvial ridges along the Mississippi 

River to –2 feet NGVD near the locally maintained levee south of Lake Pontchartrain. 

Approximately 25 percent of the developed area is below sea level. The design elevation of the 

HPS levees varies from 13 feet on the west to 12 feet on the east. There are also floodwall 

segments along the line of protection that consists of sheet-pile walls or concrete capped 

sheetpile walls constructed on the top of the levee. The line of protection was designed to 

provide protection from the Standard Project Hurricane. As designed, there is a total of 

approximately 9.5 miles of earthen levees, 1 mile of floodwall, one pump station, and five 

drainage structures, three swing gate closures for road and rail crossings, and one open gap for a 

rail crossing. The MRL is generally designed to elevation 26 feet with a 10 foot crown and a 1 V 

on 3 H slope on the land side and a 1 V on 4 H on the flood side. Similarly, the Bonnet Carre 

guide levee is generally designed to elevation 20.3 feet with a 10 foot crown and a 1 V on 3 H 

slope on the land side and a 1 V on 4 H on the flood side. 

The basin is a mix of industrial and residential areas. The area between the HPS and Lake 

Pontchartrain is essentially a wetlands area. There are two sub-basins in the basin as shown in 

Figure 1. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-2

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 1. St. Charles Basin with sub-basins and annotations for significant features 

The St. Charles HPS is made up of Reach SC1 – SC4 which begins at the Bonnet Carre 

Guide levee and extends eastward to the St. Charles – Jefferson Parrish border. 

Reach SC1 is approximately 17,000-ft-long earthen levee (with a geotextile blanket) and 

contains (1) the Bayou Trepagnier Pump Station and Drainage Structure with a transition 

sheetpile wall, (2) a pipeline crossing and (3) the Good Hope floodwall. It was designed to a net 

grade of 13 ft MSL. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-3

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Intersection of St. Charles HPS and the Bonnet Carre Guide Levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-4

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Trepagnier Drainage Structure 

Trepagnier Pump Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-5

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pipeline Crossing with sheetpile transition to levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-6

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Industrial access road closure gate 

Reach SC2 is approximately 12,000-ft-long earthen levee (with a geotextile blanket) and 

contains (1) the Cross Bayou Drainage structure and the Gulf South Pipeline crossing. It is 

designed to an elevation of 12.5 ft. There is a 500-ft transition from 12.5 to 12 ft where SC3 

begins. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-7

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Cross Bayou Drainage Structure 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Gulf South Pipeline Crossing 

Reach SC3 is approximately 24,000-ft-long earthen levee (with a geotextile blanket) and 

contains (1) The St.Rose, Almedia, and Walker Drainage structures, (2) the I310 Floodwall with 

one access gate, and (3) the railroad crossing near the airport runway extension. The RR crossing 

closure gate was not in place during Hurricane Katrina, but has since been completed. It was 

closed by sandbags for Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-9

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

St. Rose Drainage Structure 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-10

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

I-310 Floodwall 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-11

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

I-310 Floodwall access closure gate (normally closed) with 24-ft closure gate 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-12

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Alemedia Drainage Structure 

Walker Drainage Structure with sheetpile transition 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-13

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Railway Gated Crossing at Airport Levee (not constructed during Katrina) 

Railway Gated Crossing at Airport Levee (not constructed during Katrina). HPS joins Airport Levee just 

north of the Railway Gated Crossing 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-14

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach SC4 is approximately 8,048-ft-long earthen levee with most of it having an embedded 

sheetpile wall in its crown. It is designed to an elevation of 27.9 ft. It extends from where the 

HPS intersects the airport runway extension levee to the St. Charles – Jefferson Parish boundary, 

then proceeds southward to US 61, and on to the railroad crossing. Significant features are (1) an 

abrupt 3-ft drop in elevation at one 90 deg turn in the wall, (2) the 24-ft gap at the railroad 

crossing, (3) the US 61 crossing has no closure gate, and (4) the HPS ends at the railroad 

crossing, with the remainder of the parrish boundary line at the same elevation as the RR until it 

intersects the MRL. The RR crossing gap was sandbagged during Katrina. 

Floodwall as it turns east by the Airport runway extension 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-15

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Corner of sheetpile floodwall as it turns south by the Airport runway extension (3 ft drop in elevation) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-16

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Floodwall as it turns south just north of the railroad crossing 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-17

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Railroad crossing has no closure gate. During Katrina it was sandbagged. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-18

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

St. Charles – Jefferson Parish Boundary Levee ends North of US 61 Highway – gap at this crossing 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-19

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

St. Charles – Jefferson Parish Boundary Levee South of US 61 Highway – no closure at this crossing 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-20

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Short Transition Section of HPS from South of US 61 toward railroad 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

St. Charles –Jefferson Levee at South side of US 61 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-22

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The basin protection then continues westward as the MRL at a design elevation of 26 ft. The 

MRL is an earthen levee with a 10-ft crown. No major structure or pipeline passes through the 

MRL. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-23

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Mississippi River Levee at the St. Charles-Jefferson boundary at Hwy 48 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-24

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Top of Mississippi River Levee at the St. Charles-Jefferson boundary 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-25

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

I-310 passing over the Mississippi River Levee 

At the east extend of the basin, the MRL intersects the Bonnet Carre guide levee, which 

continues the protection northward on the west side of the basin until it reached the HPS. This 

stretch contains the intersection with the spillway and US 61. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Mississippi River Levee intersection with Bonnet Carre Guide Levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Bonnet Carre Spillway intersection with guide levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-4-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Bonnet Carre Guide Levee Looking North from the Spillway gates 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-1

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 5 

Plaquemines Basin 

Background 

The Plaquemines (PL) Basin is made up of 11 sub-basins as shown in Figure 1. It is 

generally composed of the Plaquemines Parrish along the east and west banks of the Mississippi 

River south of Mile 82. Both the west and east bank protection includes the Mississippi River 

levees as a part of each sub-basin. There are 134 miles of MRL and floodwall, 53 miles of 

hurricane protection, 12 miles of floodwall, 19 pump stations, a 110-foot small boat lock, and a 

marine floodgate. The damage consisted of 20 miles of MRL and HPS levee, 9.4 miles of 

floodwall, and five pump stations. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-2

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 1. Plaquemines Parrish with levee protection footprint and sub-basins (PL1 – PL11) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-3

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

PL 11 

PL 11 begins along the west back of the MRL, shown in Figure 2, is generally bounded on 

the east by the Mississippi River, the Intercoastal Waterway on the west, the PlaqueminesOrleans Parrish boundary on the north, and the Hero Canal on the south. 

Figure 2. PL 11 and PL 1 Sub-Basins with Reach Beginning and Ending Points (Red Dots) 

The federal levee begins at the MRL inside the U.S. Coast Guard station. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-4

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-5

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-6

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-7

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Back levee continues outside U.S. Coast Guard toward GIWW Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Drainage structure under Back Levee outside U.S. Coast Guard station. Screw gate closure on culvert 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-9

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Back Levee intersects Hwy 406 looking toward the GIWW and General De Gaulle Bride overpass. 

No closure at Hwy 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-10

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Back Levee passed under General De Gaulle onramp from Hwy 406. Buckling of concrete slab on levee 

under ramp. 

The interior Orleans-Plaquemines Parrish levee ends at the GIWW. The Federal Back levee 

along the GIWW the proceeds south and passes under the General De Gaulle Highway bridge 

overpass. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-11

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-12

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Going south along the GIWW to the Belle Chase Pump Station. 

Sub-basin continues south along the GIWW to the Highway 23 and Railroad Bridge Crossing 

and Tunnel under GIWW. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-13

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Highway 23 Bridge over GIWW and Levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-14

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-15

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Highway 23 Tunnel under the GIWW 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-16

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pipe crossing over the GIWW levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-17

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

A point of levee erosion along the GIWW, however most is in good condition 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-18

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical levee section. Numerous gates across levee to contain cattle 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-19

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical levee section along the GIWW 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-20

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Plaquemines Pump Stations 1 and 2 

Plaquemines Pump Stations 1 and 2 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

PL 11 Back levee intersects Hwy 23 north of Jesuit Bend 

The MRL then forms the remaining section of PL 11 as it goes north along the river to the 

point where it intersects the Plaquemines Parish interior levee inside the U.S. Coast Guard 

station. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-22

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The crown of the levee is generally 10 feet wide, with most paved and some stretches gravel. Numerous 

pipes cross over the MRL in this area 

A number of off-load facilities are located along this portion of the MRL, similar to this grain loading facility 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-23

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

There were a few cases of erosion along the MRL 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-24

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

More typical condition along this section of the MRL, with concrete paved or stone crown and concrete 

armored floodside 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-25

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

PL 2 

PL 2 begins at this point and continues south toward the Alliance Refinery as a non-federal 

levee. Figure 3 shows the sub-basins PL2, PL3 and PL 6. 

Figure 3. Location of sub-basins PL 2, PL 3, and PL 6 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Ollie Pump Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Ollie Pump Station 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The condition of the non-federal levee varies. The crown is generally in better condition if 

cows are present on the levee. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The PL 2 back levee comes to an end back at Hwy 23 just north of the Alliance Refinery. It 

then is completed by the MRL as it goes north along the river. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-30

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

PL 3 

PL 3 then begins at this point, just north of the Alliance Refinery, as a non-federal levee and 

proceeds southward to Myrtle Grove, where it intersects Hwy 23. Levee conditions are generally 

poorer than the PL 2 back levee, with large overgrowth being common. Many places are 

impassable by 4 WD vehicles. 

The condition of the non-federal levee also varies. The crown is generally poor condition, 

with much impassable by motorized vehicle. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-31

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Continuing southward along the MRL, the sub-basin enclosed by the Citrus Lands Back 

Levee will not be included. There was an extensive breach in this levee and repairs were 

difficult. Below is the initial repair. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-32

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Citrus Lands Back Levee repairs to breach 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-33

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Citrus Lands Back Levee repairs to breach 

The Back Levee crosses Hwy 23 and connects to the MRL. The sub-basin enclosure then 

goes north as the MRL. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The Back Levee intersects the MRL after it crosses Hwy 23 

As the MRL proceeds north, numerous locations of erosion occurred due to overtopping. 

Some were repaired and some were not extensive enough to require repair. Debris is located on 

the levee top and slopes. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-35

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-36

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-37

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-38

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Also the MRL has numerous locations where the concrete armor was eroded and was 

replaced with stone. More cases of erosion. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-39

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-40

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

PL 4, 5, 7, 8, 9, and 10 

These sub-basins are on the west bank, south of St. Jude. These sub-basins are in a project 

named New Orleans to Venice (NOV) Hurricane Protection Project. 

NOV 

The Mississippi River Levees also serve as the hurricane protection system south of St. Jude 

and is part of the NOV Hurricane Protection Project. On the East Bank, the project extends 16 

miles from Phoenix down to Bohemia. On the West Bank it extends 37 miles from St. Jude to 

Venice.

Figures 4a and 4b give a comparison the NOV project definitions, and the sub-basin PL 4, 5, 

7, 8, 9, and 10 locations. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-41

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 4a. New Orleans to Venice, LA, Hurricane Protection Project 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-42

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 4b. Sub-basins PL 4, 5, 7, 8, 9, and 10 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-43

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

West Bank Back Levees 

St. Jude to City Price 

The St. Jude to City Price reach includes 3 miles of enlarged back levees from St. Jude to 

City Price (between approximate river miles 47.1 and 43.9 Above Head of Passes (AHP)). This 

levee was constructed to elevation 7.0 feet NGVD as a non-federal levee. The non-federal levee 

was later incorporated into the federal project at 12.5 ft (NGVD). The NOV project area includes 

approximately 15,600 acres of land including 4,300 acres in Reach A; 3,800 acres in Reach B-1; 

2,300 acres in Reach B-2; 4,500 acres in Reach C and 700 acres in the St. Jude to City Price 

area. 

Reach A includes 13 miles of enlarged back levees from City Price to Tropical Bend and two 

54-in. flap-gated culverts (between approximate river miles 43.9 and 30.7 AHP). It consists of 

approximately 12.8 miles of levee system with a net elevation of 12.5-14.5 feet and includes 

floodwalls at the Hayes Canal and Gainard Woods Pump Stations. The levee enlargement 

consisted of a marsh side embankment with a wave berm. The base of the levee incorporated 

geotextile fabric with a sand blanket and a clay cap at least 2 feet thick. The embankment was 

constructed of uncompacted clay, though a sand core may have been substituted in places. The 

main levee cross section is 1V on 3H.

Reach B-1 includes 12 miles of enlarged back levees with a net elevation of 15 ft from 

Tropical Bend to Fort Jackson (between approximate river miles 30.7 and 20.5 AHP) and a 

marine floodgate at Empire. The main levee cross section is 1V on 4H. The Flood Side (FS) and 

Protected Side (PS) berms generally vary from 1V on 15-20H and from 1V on 12-20H, 

respectively. The reach also includes a flood gate at Empire and floodwalls at the Bayou Grand 

Laird (I- and T-wall) and Sunrise Pump Stations. The Empire Floodgate is in the Empire to Gulf 

Waterway and consists of a reinforced concrete U-shaped gated bay with a steel gate hinged at 

the bottom, guide walls and fenders, inverted T-wall reinforce concrete floodwalls extending 

about 150 feet on each side of the structure, access road and breakwater. 

Reach B-2 includes 9 miles of enlarged back levees with a net elevation of 15 ft from Fort 

Jackson to Venice between approximate river miles 20.5 and 10.4 AHP and includes floodwalls 

at the Venice Pump Station. The levee consists of a sand core with hydraulic clay fill. The levee 

construction occurred in three stages or lifts. The main levee cross section is 1V on 4H with the 

Flood Side (FS) and Protected Side (PS) berms generally varying from 1V on 15-74H FS and 

from 1V on 29-71H PS.

West Bank River Levee 

The West Bank River Levee (WBRL) includes 34 miles of West Bank Mississippi River 

levees built to a net design elevation of 16 to 17 ft, from City Price to Venice (between 

Mississippi river miles 44 to 10 AHP) (Note: the lock at Empire is a State of Louisiana facility.)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-44

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

PL 7 

Sub-basin PL 7 begins near City Price at Diamond Pump Station and continues past Hayes 

Pump Station and on southward near Homeplace as shown in Figure 4b. 

PL 8 

Sub-basin PL 8 begins at this location and proceeds on to near Empire Lock. 

PL 9 

Sub-basin PL 9 begins near Empire Lock and proceeds on to near Sunrise Pump Stations. 

PL 10 

Sub-basin PL 10 begins at Sunrise Pump Stations and proceeds on to near Venice. 

East Bank 

PL 4 and PL 5 

The back levee begins at Phoenix and proceeds southward to Bohemia. The separation 

between PL 4 and PL 5 occurs near Pointe a La Hache. Reach C of the NOV and the East Bank 

of the MRL enclose these sub-basins. Reach C consists of approximately 16 miles of enlarged 

back levees with a net elevation of 17 feet. The back levee has a sand core with clay blanket. It 

was enlarged with hauled fill and raised from approximately 14-feet elevation to the 17-foot 

design level. It includes floodwalls (I-type sheet piling) at the pump stations near Bellevue and 

Pointe a La Hache. Construction of the levee to date has included three of the designed four lifts. 

It lays between approximate river miles 59.3 and 44.3 AHP and 10 flap-gated culverts. 

The NOV was damaged by Hurricane Katrina when it made landfall near Buras-Triumph, 

which is on Reach B-1. The storm produced storm surge levels that exceed the level of the 

constructed protection. Numerous breaches occurred along the back levees on both the east and 

west bank sides of the NOV project. Levees were overtopped and breached, resulting in 

extensive erosion and scour, along both the back levees and the Mississippi River levees (as 

enlarged for hurricane protection). In addition there was damage to the floodgate at Empire and 

to the floodwalls along the MRL and back levees.

PL 1 and PL 6 

The final two sub-basins (PL 1 and PL 6) are on the east bank of the river across from PL 11 

(Figures 2 and 3). The non-federal back levee begins near the Plaquemines-St. Benard Parrish 

boundary and continues south, ending south of Belair. The protection level is at elevation 6 feet. 

These sub-basins are closed by the MRL as it proceeds north along the river to the parrish 

boundary. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-45

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Risk Model Idealization 

The Plaquemines Parrish Basin was descretized into 11 sub-basins (PL1 – PL11) as shown in 

Figure 5. The sub-basins were defined to correspond to the known interior drainage areas. This 

reach idealization follows from the basin description information presented previously, which 

was collected from project documents and field inspections. Figures 6, 7 and 8 show the 

elevations for the Plaquemines Parrish HPS: Pre-Katrina-at the time of Katrina, Current-as of 

1 Jun 2007, and the Authorized-at the time Katrina occurred. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-46

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 5. Plaquemines Parrish Basin reaches (PL01-PL28) and sub-basins (PL 1 - PL 11) definition for 

use in the risk model 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-47

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 6. Elevations for the Plaquemines Parrish Basin for the Pre-Katrina HPS (in place when Katrina 

occurred) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-48

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 7. Elevations for the Plaquemines Parrish Basin for the Current HPS (as of 1 June 2007) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-49

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 8. Elevations for the Plaquemines Parrish Basin for the Authorized HPS (authorized at the time of 

Katrina) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-50

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Plaquemines Transitions 

Feature 

Number 

Transition 

Type Reach 

Width 

(ft) 

Elevation (ft) 

(NAVD88 2004.65) Sub-Basin Description of Feature 

132 P PL1 400.0 8 PL12 Pump Station - Belle Chase #2 

133 P PL1 175.0 10 PL12 Pump Station - Belle Chase #1 

134 P PL4 280.0 10 PL2 Pump Station - Upper Ollie 

135 P PL6 100.0 8 PL3 Pump Station - Wilkerson Canal 

136 P PL8 120.0 18 PL8 Pointe A Lache West Pump 

Station 

137 P PL8 170.0 10 PL7 Diamond Pump Station 

138 P PL10 342.0 16 PL11 Hayes Pumping Station 

139 P PL11 550.0 17 PL8 Gainard Woods Pump Station 

140 P PL16 1010.0 20 PL9 Sunrise Pumping Station 

141 G PL16 635.0 19.5 PL9 Empire Flood Gate 

142 P PL17 627.0 19 PL10 Venice Pumping Station 

143 P PL17 975.0 19 PL10 Grand Liard (Buras) PS 

144 P PL22 100 18 PL4 Bellevue Pumping Station 

145 P PL27 175 20 PL5 Pointe A La Hache East 

Pumping Sation 

146 P PL25 100 10.5 PL6 Belair Pump Station 

147 P PL25 200 14 PL1 Scarsdale Pump Station 

148 P PL24 80 9 PL1 Braitwaithe Pump Station 

149 U PL2 2200 8 PL2 Unprotected area between PL11 

and PL2 

150 U PL1 1650 5 PL1 Unprotected area between 

PLAQ and STB 

151 U PL1 1730 18 PL1 MRT between PLAQ and STB 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-5-51

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Plaquemines Reaches 

Reach 

No. 

Length 

(ft) 

Pre-Katrina Elevation (NAVD88 

2004.65) 

Reach Type 

(1) 

Foundation Material Type (H, 

C, P) (2) 

Subbasin 

Reference (3) 

81 22,000 6.00 L H PL11 

82 41,525 8.50 L H PL11 

83 57,470 18.10 L C PL11 

84 50,610 8.50 L H PL2 

85 36,605 16.40 L C PL2 

86 60,615 6.40 L H PL3 

87 25,865 15.70 L C PL3 

88 17,170 11.20 L H PL7 

89 39,195 16.20 L C PL7 

90 27,100 13.50 L H PL7 

91 19,120 13.60 L H PL8 

92 13,774 12.70 L H PL8 

93 6,635 13.80 L H PL8 

94 49,470 16.30 L C PL8 

95 6,160 14.90 L H PL8 

96 26,710 15.00 L H PL9 

97 78,500 14.70 L H PL10 

98 79,100 15.00 L C PL10 

99 22,740 13.90 L C PL9 

100 51,200 16.60 L C PL4 

101 32,235 15.60 L C PL5 

102 50,475 17.30 L H PL4 

103 29,050 17.50 L H PL5 

104 62,810 12.00 L H PL1 

105 30,940 12.40 L H PL6 

106 61,710 18.60 L C PL1 

107 25,225 17.00 L C PL6 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-1

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 6 

St. Bernard Basin 

The St. Bernard (STB) basin is defined by the protection system along the GIWW to the 

north, MRGO to the east, Caernarvon Canal to the south, and the Mississippi River and Inner 

Harbor Navigation Canal (IHNC) to the west. Like New Orleans East, it is essentially separated 

into two distinct areas, a residential/commercial area on the south side of the basin and a 

marshlands area on the north side. These two sections are separated by a non-federal, interior 

local levee that runs across the basin from the northwest to the southeast. This area, along with 

other pertinent information relative to the IPET assessment, is depicted in Figure STB-1. 

The levee and floodwall system surrounding the STB basin consists of approximately 

157,800 linear feet of varying levels of protection. This provides protection for an area of 

approximately 81 square miles for the entire basin. The residential area makes up approximately 

27 square miles of the basin. In addition, there are two water control sector gate structures along 

the MRGO at Bayou Bienvenue and Bayou Dupre. There are a total of nine pump stations within 

the basin, primarily along the interior local levee; one of these is located along the Caernarvon 

Canal. These major structures are also depicted in Figure STB-1. 

Like the other basins, the STB was constructed during different times and modified at 

various places since the last Design Memorandums. For the purposes of IPET and coupled with 

varying versions of the most recently completed Design Memorandums (DM), the STB basin is 

separated into three major stretches. These are as follows: 

• North Side of IHNC Lock thru Caernarvon Canal. This stretch of the STB basin 

represents the exterior hurricane protection system and begins at the tie-in to the north 

side of the IHNC Lock, continues northeast along the GIWW, turns and follows the 

MRGO to the southeast, then goes west back towards the tie-in to the Mississippi River 

levee. 

• Mississippi River Levee (MRL). This section is the flood protection system along the 

Mississippi River system that contains a combination of floodwalls and levees. For 

numbering purposes, the MRL reaches begin at the tie-in with the Caernarvon Canal and 

runs northwest along the Mississippi River until it ties back in with the south side of the 

IHNC Lock. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-2

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

• Interior Local Levee (ILL). This section of levee and floodwall separates the residential 

and marshland areas of the STB basin. The ILL is actually owned by the State of 

Louisiana and maintained by the LADOTDD and Lake Borgne Basin Levee District. 

USACE was provided a one-time waiver from policy and was tasked with repairing the 

damage to this levee following Katrina. The ILL basically splits the basin in two and 

begins along the IHNC and heads generally in a southeast direction along the middle of 

the basin. The ILL “wraps” around the Violet Canal and goes generally east before 

turning south and tying back into the exterior levee protection system along the 

Caernarvon Canal. 

Figure STB 1. St. Bernard basin 

STB Basin – Layout of Reaches for Risk Model by Physical 

Feature (Pre-Katrina) 

Within these major stretches of the STB basin there are shorter reaches, which are defined by 

physical changes in the protection system, i.e. switching from floodwall to levee, etc., or by 

significant changes in geotechnical parameters. Within each reach, there are specific “key 

points” whose reliability needs to be determined in order to calculate the effect on the overall 

reach being evaluated. An example of a “key point” would be a closure gate at a road or rail line 

Lake Borgne MRGO

Violet Canal

4

7

3

GIWW

Mississippi River

IHNC

1/6

2

5

Paris Road

Caernarvon Canal

Interior Local Levee

Hurricane Protection

Levee and Floodwall

# Pump Station

Miss. River Levee

Parish Boundary

Bayou Dupre

Control Structure

Bayou Bienvenue

Control Structure

5

8
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-3

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

crossing along a floodwall. Task 10 engineers reviewed existing plans, damage survey reports, 

and conducted field verification inspections to ensure each basin was accurately defined within 

the system. As a part of the field verification inspections, GPS coordinates were obtained and 

stationing from DM’s and “as-built” plans was verified. For each basin, this information was 

transformed into a spread sheet and then a system map for each basin, as shown in 

Figure STB 2A. Finally, digital photographs with incorporated notes were developed to 

compliment the spread sheets and system map for further clarification. This collection of 

information was then categorized to get a clear picture of how the basin should be defined for 

risk assessment purposes. A summary of the reach and point definitions for STB is shown in 

Figure 2A with a brief supporting narrative on each reach. The layout shown in Figure STB 2 

and the narrative that goes along with this figure relates to the pre-Katrina condition. Task Force 

Guardian (TFG) made several improvements to the levee/floodwall system which changes the 

risk for various reaches. These changes by reach are detailed in the next section. 

Task 10 basin reach definitions for STB start at the tie-in to the north side of the IHNC Lock. 

The numbering system for reach definitions continues along the exterior protection system along 

the GIWW, MRGO, Caernarvon Canal, and then up through the Mississippi River levee system 

to where it ties back into the south side of the IHNC Lock. Finally, the interior local levee 

reaches numbers begin at the IHNC and extend until it ends at the Caernarvon Canal. Please 

refer to Figure STB1, STB2A, and STB2B for further clarification. The STB basin is 

summarized by the reaches as follows: 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-4

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 2A. St. Bernard Basin – Reaches Defined 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-5

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 2B. St. Bernard Basin Reaches – Insets 1-6 

Reach STB1 (INHC to Caernarvon) 

This reach is defined by 1,427 linear feet of concrete capped I-wall that ties into the northeast 

side of the INHC Lock and generally follows the IHNC north, see Figure STB 3, which shows 

the beginning of the reach where it ties into the northeast lock wall of the IHNC. There is one 

key point (stb1a) at the beginning of this reach where the natural ground line transitions between 

the lock wall and I-wall, as shown in Figure STB 4. The reach ends at the south end of the I-wall 

failure that occurred in the Lower 9th Ward. Other sections of this reach were overtopped during 

Katrina, but did not fail, in particular near the Claiborne Bridge, as shown in Figure STB 4. The 

approximate weighted average top of wall elevation for this reach was 13.0 (NAVD88) prior to 

Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-6

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Figure STB 3. I-wall section tying into IHNC lock wall (standing on top of lock wall and looking away from 

IHNC) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-7

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 4. Section of Reach STB1 overtopped during Katrina (location is just north of Claiborne 

Avenue Bridge) 

Reach STB2 (IHNC to Caernarvon) 

This reach was defined by a 4,038-ft stretch of I-wall between the Claiborne Avenue Bridge 

and the railroad bridge near Florida Avenue. There were two separate breaches of this I-wall 

during Katrina. The southern section of I-wall that failed in the Lower 9th Ward area of 

St. Bernard during Hurricane Katrina is shown in Figure STB 5. Figure STB 6 shows the 

authorized design section for this wall, but note the actual elevations for the top of wall were 

closer to 13.0 when referencing NAVD88 (2004.65) datum. The northern section of I-wall that 

failed was near the blue railroad bridge close to Florida Avenue. This breach is depicted in 

Figure STB 7. There were no “key points” within this entire reach. Refer to changes made by 

Task Force Guardian (TFG) in the post-Katrina narrative. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 5. Failed I-wall section of Lower 9th Ward (Reach STB2) (looking south along IHNC) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-9

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 6. Pre-Katrina I-wall design section for Reach STB2 (depicts authorized elevations, not actual 

elevations) 

Figure STB 7. North I-wall failure within Lower 9th Ward (Reach STB2) (Note: pipes on right side of photo 

are associated with Pump Station #5) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-10

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB3 (INHC to Caernarvon) 

This reach is defined by an existing T-wall that is located adjacent to the Surekote Road 

ramp over the floodwall. The reach is approximately 807 feet in length. There are multiple key 

points within this reach including the closure gate E-1, as shown in Figure STB 8. Closure gate 

S-1 at Harbor Road and Florida Avenue and railroad closure gate E-2 are also key points within 

this reach. The approximate weighted average height of this wall prior to Katrina was elevation 

12.5 (NAVD88). 

Figure STB 8. Existing T-wall near Florida Ave. Bridge (Reach STB3) (looking from protected side towards 

IHNC) 

Reach STB4 (INHC to Caernarvon) 

This reach is defined by 726 linear feet of I-wall that ties into the levee along the GIWW. 

There are no key points in this reach. This reach ends where the I-wall ties into the levee. This 

section of I-wall was overtopped during Katrina, but suffered no major damage. It had an 

approximate weighted average top of wall elevation of 13.3 (NAVD88) prior to Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-11

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB5 (INHC to Caernarvon) 

This reach is defined by 890 linear of levee between the end of the IHNC I-wall and the 

floodwalls surrounding closure gates S-2 and S-3 near the Southern Scrap Building. There are no 

key points within this reach. This levee is depicted in Figure STB 9. There was no significant 

damage to this section from Katrina. It had an approximate weighted average elevation of 13.1 

(NAVD88) prior to Katrina. 

Figure STB 9. Beginning of Reach STB5 at end of I-wall (Note TFG improvements to Reach STB4 in 

foreground) 

Reach STB6 (IHNC to Caernarvon) 

This reach is defined by a combination of I-walls and T-walls surrounding closure gates S-2 

and S-3. The total length of the reach is 340 feet with T-walls located around the gate closures. 

These walls were overtopped during Katrina but did not fail. The approximate weighted average 

top elevation of this reach was 13.0 (NAVD88) prior to Katrina. Reference improvements to the 

scour protection made by TFG in the post-Katrina reach description. The gates themselves serve 

as the two “key points” within this reach, stb6a and stb6b, respectively. 

Reach STB7 (IHNC to Caernarvon) 

This reach is defined by 25,722 feet of levee. There is one pipe crossing located within this 

reach but it does not represent a significant departure from the levee section and can be ignored 

for the purposes of this assessment. There was no significant damage to this section of levee 

from Katrina. The approximate top of levee varied along this reach between 13.5 to 16.5 

(NAVD88), but had a weighted average elevation of 15.1 (NAVD88) prior to Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-12

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB8 (IHNC to Caernarvon) 

This reach is defined by a 1,016-ft stretch of concrete capped I-walls and T-walls just west of 

the Paris Road overpass along the GIWW. It is often referred to as the Paris Road floodwall. 

There is one key point (stb8a), a gate closure, within this floodwall reach. This structure was 

overtopped during Katrina, but only suffered scour damage. See Figure STB 10 for a photograph 

depicting this wall during repairs following Katrina. The approximate weighted average top of 

wall elevation for this reach was 12.4 (NAVD88) prior to Katrina. 

Figure STB 10. Paris Road Floodwall (Reach STB8) (looking east from the west end of the floodwall) 

STB9 (IHNC to Caernarvon) 

This reach is defined by 7,260 linear feet of levee. It goes between the Paris Road floodwall 

and the floodwalls just prior to reaching the Bayou Bienvenue control structure. There are no key 

points within this reach. There was some overtopping in this reach during Katrina, but no major 

damage to the levee. The approximate weighted average for this reach was elevation 17.9 

(NAVD88) prior to Katrina. 

STB10 (IHNC to Caernarvon) 

This section is defined by a relatively short 229-ft stretch of floodwall located just northwest 

of the Bayou Bienvenue Control Structure. It consists of a 61-ft section of I-wall at elevation 

17.0 (NAVD88), a 107-ft stretch of T-wall at elevation 15.5 (NAVD88), and another 61 feet of 

I-wall at elevation 17.0 (NAVD88). The weighted average top of wall across the entire reach is 

16.3 (NAVD88). During original construction, this short section of wall was to serve as an 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-13

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access point for an industry that was to be located near the Bayou Bienvenue Control Structure. 

However, the industry went out of business, and it has never been utilized, and the flood gate 

remains permanently closed. This section of wall was overtopped during Katrina as evidenced in 

Figures STB 11 and STB 12. Since the gate remains permanently closed, there are no “key 

points” within this reach. 

STB11 (IHNC to Caernarvon) 

This reach is defined by a short section of levee between the floodwall described in reach 

STB10 and the beginning of the floodwall leading to Bayou Bienvenue Control Structure. The 

levee section is only 96 feet long and received some damage during Katrina at the transitions to 

adjoining I-wall sections, as shown in Figure STB 12. The levee had an average weighted 

elevation of 16.0 (NAVD88) prior to Katrina. There are no “key points” within this short reach. 

Figure STB 11. Floodwall Northwest of the Bienvenue Structure (Reach STB10) (looking southeast toward 

Bayou Bienvenue just after Katrina) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-14

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 12. Short levee section between floodwalls (Reach STB11) 

Reach STB12 (IHNC to Caernarvon) 

This reach is defined by the Bayou Bienvenue Control Structure and surrounding floodwalls 

on either side of it. Prior to Katrina, this reach was made up of 77 feet of concrete capped I-wall 

at elevation 17.0 (NAVD88), 187 feet of T-wall surrounding the control structure itself at 

elevation 15.0 (NAVD88), and 1,208 feet of uncapped, sheetpile I-wall at elevation 18.5 

(NGVD). As shown in Figures STB12, 13, and 14, the surrounding floodwalls received heavy 

damage and the uncapped I-wall section breached during Katrina. The sector gate closure itself 

serves as the only “key point” (stb12a) within this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-15

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 13. Uncapped I-wall Section on Southeast side of Bayou Bienvenue Control Structure (looking 

southeast along the MRGO) 

Figure STB 14. Picture of breached Bienvenue I-wall from Protected Side 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-16

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB13 (IHNC to Caernarvon) 

This reach is defined by a 19,858 linear feet stretch of levee. Prior to Katrina, it began at the 

end of the uncapped sheetpile I-wall on the southeast side of the Bayou Bienvenue Control 

Structure and continued up to the beginning of an embedded, uncapped I-wall section along the 

MRGO. This section of levee was overtopped during Katrina and was heavily damaged. It had 

an approximate weighted average top elevation of 17.5 (NGVD) based upon a 1998 survey. 

There was one section of embedded uncapped, sheetpile I-wall in about the middle of this reach, 

but the top of the wall essentially was at the top of the levee. Due to this fact, this section of 

I-wall will not be included separately in the analysis and is included as part of the overall reach 

characteristics. 

Reach STB14 (IHNC to Caernarvon) 

This reach was defined by a 2,427-ft section of uncapped I-wall embedded within the levee. 

The sheetpile I-wall was installed in 1992 as part of USACE repairs along the MRGO. This 

section of wall was heavily damaged during Katrina, as evidenced in Figure STB 15. The 

weighted average top elevation of the uncapped I-wall was 18.5 (NGVD) prior to Katrina, and it 

had a free standing height of approximately 3.5 feet. As shown in the photo, there are pipe 

crossings along this reach, but they extend over the levee and do not represent a major change in 

design or performance parameters; therefore, they are ignored in the risk assessment for this 

purpose. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-17

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 15. Breached I-wall along MRGO 

Reach STB15 (IHNC to Caernarvon) 

This reach consists of a 1,745-linear-ft stretch of levee between uncapped, sheetpile I-wall 

reaches. This section was overtopped and heavily damaged during Katrina. The weighted 

average height of this levee section was approximately 16.4 (NGVD) prior to Katrina based 

upon a detailed 1998 survey. There are no “key points” within this reach. 

Reach STB16 (IHNC to Caernarvon) 

This reach consisted of a 2,560-ft stretch of uncapped, sheetpile I-wall that was installed in 

1992 as part of USACE repairs along the MRGO. This section also was overtopped and heavily 

damaged during Katrina. The weighted average top of uncapped I-wall along this reach was 

elevation 18.5 (NGVD) prior to Katrina. It had an approximate free standing height of 3.5 feet. 

There are no “key points” within this reach. 

Reach STB17 (IHNC to Caernarvon) 

This reach consists of 566 feet on levee between uncapped, sheetpile I-wall reaches. The 

approximate weighted average top of levee elevation was 16.5 (NGVD) prior to Katrina based 

upon a detailed 1998 survey. This section of levee was overtopped during Katrina and heavily 

damaged. There are no “key points” within this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-18

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB18 (IHNC to Caernarvon) 

This reach consists of 359 feet of uncapped, sheetpile I-wall installed in 1992 as part of 

USACE repairs along the MRGO. The wall had an approximate top elevation of 18.5 (NGVD) 

with a free standing height of roughly 3.5 feet. It was overtopped and heavily damaged during 

Katrina. There are no “key points” within this reach. 

Reach STB19 (IHNC to Caernarvon) 

This stretch of levee runs from the end of the uncapped I-wall in Reach STB18 to the 

northwest transition wall leading to the Bayou Dupre Control Structure. It is approximately 

4,994 linear feet and there are no “key points” within this length. It had an approximate weighted 

average top elevation of 18.7 (NGVD) prior to Katrina. This reach was heavily damaged during 

Katrina. 

Reach STB20 (IHNC to Caernarvon) 

This reach includes the Bayou Dupree Control Structure and the adjoining transition flood 

walls. Prior to Katrina, the reach started with 92 linear feet of precast concrete sheetpile wall on 

the northwest side of the gate structure, 69 feet of T-wall on the northwest side of the gate 

closure, 134 feet across the closure structure, 69 feet of T-wall on the southeast side of the 

closure structure, and 92 feet of precast concrete sheetpile wall on the southeast side of the 

closure structure. While the concrete sheetpile walls were designed at a higher elevation, 

settlement across this area left all walls roughly at elevation 15.2 (NAVD88) prior to Katrina. As 

shown in Figure STB 16, the northwest precast concrete sheetpile I-wall failed during Katrina. 

The only “key point” in this reach is the closure structure itself. 

Reach STB21 (IHNC to Caernarvon) 

This reach starts with the levee tie-in to the southeast side of the Bayou Dupre Control 

Structure adjoining floodwall. The approximate length of this reach is 25,562 linear feet, and it 

has no “key points” within the reach. Prior to Katrina, it had a weighted average top elevation of 

19.1 (NGVD). The reach was heavily damaged during Katrina from overtopping. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-19

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Figure STB 16. Damage at Bayou Dupre Control Structure (note breached section of precast concrete pile 

wall) 

Reach STB22 (IHNC to Caernarvon) 

This reach was defined by a 1,401-ft section of uncapped, sheetpile I-wall that was installed 

during 1992 USACE repairs along the MRGO. The wall had a top elevation of 19.0 (NGVD) 

prior to Katrina and a free standing height of approximately 4.0 feet. There were no “key points” 

within this reach. There was scour damage in this area as a result of Katrina. 

Reach STB23a (IHNC to Caernarvon) 

This reach is defined by the remaining levee along the MRGO between the end of Reach 

STB22 and where it turns away from along the Caernarvon Canal. The weighted average top of 

levee elevation for this reach was approximately 19.5 (NGVD) prior to Katrina. There are no 

“key points” within this reach. Sections of this reach did receive damage during Katrina. 

Reach STB23b (IHNC to Caernarvon) 

This reach consists of 5,709 feet of levee that begins at the MRGO and continues to the 

intersection point with the interior local levee. Although the levee along the Caernarvon Canal 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-20

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continues, a new reach had to be defined because of the potential for varying consequences due 

to the presence of the interior local levee. This area of levee received minor damage during 

Katrina. There are two basic areas along this reach where repairs are different. The weighted 

average top elevation of this reach was approximately 17.0 (NAVD88) prior to Katrina. There is 

one “key point” within this reach (stb23bd), which is the Bayou Road closure gate. 

Reach STB23c (IHNC to Caernarvon) 

This section of levee is 2,692 feet long and goes from the interior local levee to the beginning 

of the east side floodwall surrounding Pump Station #8. There was minimal scour damage along 

this stretch during Katrina. It had an approximate weighted average top elevation of 16.0 

(NAVD88) prior to Katrina. There are no “key points” within this reach. 

Reach STB23d (IHNC to Caernarvon) 

This reach consists of 231 feet of flood wall around the discharge pipes for Pump Station #8. 

There are roughly 45 feet of I-walls on both sides of the middle T-wall for the discharge pipes. 

The T-wall length is approximately 141 feet long. The top of the I-walls are elevation 17.0 

(NAVD88) and the T-wall is at 16.5 (NAVD88). The weighted average top of wall elevation is 

16.7 (NAVD88). This wall was not damaged or overtopped during Katrina. There are no “key 

points” within this reach. For a view of this reach, please refer to Figure STB 17. 

Figure STB 17. Floodwalls surrounding Pump Station #8 pipes (viewed Looking West Along Caernarvon 

Canal) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB24 (IHNC to Caernarvon) 

This reach is defined by 36,610 linear feet of levee along the Caernarvon Canal. It starts on 

the east side at the where it ties into the west end of the Pump Station #8 floodwall and continues 

until it ties into a sheetpile wall near the Caernarvon Freshwater Diversion Structure. This 

section received little or no damage during Katrina. There are several utility crossings along this 

reach, but none significant enough to warrant as a “key point.” The approximate weighted 

average top elevation of the levee is 15.4 (NAVD88) across this reach. 

Reach STB25 (IHNC to Caernarvon) 

This reach consists of 693 linear feet of uncapped, sheetpile I-wall near the Caernarvon 

Freshwater Diversion Structure. There are no “key points” within this reach. It received little or 

no damage from Katrina. The reach has a weighted average top elevation of 12.8 (NAVD88) 

taken from a physical survey following Katrina. See Figure STB 18 for a photograph of this 

reach. 

Figure STB 18. Uncapped Sheetpile I-wall near Caernarvon Canal (building on left side is on flood side) 

Reach STB26 (IHNC to Caernarvon) 

This reach consists of concrete capped I-walls that extend from the end of the sheetpile wall 

in Reach STB25 to the where the Caernarvon section ties into the Mississippi River levee. This 

section of wall is 1,104 feet long with a weighted average elevation of 13.0 (NAVD88). It was 

not damaged during Katrina. There are two “key” points within this reach, and both are closure 
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gates; one for a rail line and the other for Highway 39. The location where the floodwall ties into 

the higher Mississippi River levee is shown in Figure STB 19. 

Reach STB27 (Mississippi River Levee) 

This reach is the most southern section of levee within the STB basin, as reflected as the 

higher ground area of Figure STB 19. Refer to Figure STB 2 for reference to the reach location 

relative to the entire basin. The reach consists of 49,877 linear feet of levee, and it ends at the 

southern end of the concrete capped I-wall near the Battlefield site along the Mississippi River. 

It has a weighted average top elevation of 20.1 (NAVD88). 

Reach STB28 (Mississippi River Levee) 

This reach consists of 2,724 feet of concrete capped I-wall near the Battlefield site. There is 

one key point with this reach, and it is a small access closure gate. A typical stretch of this wall 

along with the access closure gate is shown in Figure STB 20. The weighted average elevation of 

the top of the wall for this reach is 17.6 (NAVD88). This reach was not damaged during Katrina. 

Figure STB 19. Caernarvon Canal Floodwall tie-in to higher MRL (gate is for Highway 39 closure) 
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Figure STB 20. Concrete capped I-wall along Miss. River near Battlefield (access closure gate is key point 

within this reach) 

Reach STB29 (Mississippi River Levee) 

This reach consists of 3.729 feet of levee along the Mississippi River in and around the 

Rodriguez Canal. There are four key points within this reach, all gated closures near the 

Rodriguez Canal. The reach ends where the levee ties into a floodwall near the Domino Sugar 

Plant. The weighted average top elevation of the levee along this reach is 20.9 (NAVD88). This 

reach was not damaged during Katrina. 

Reach STB30 (Mississippi River Levee) 

This reach is defined by 5,180 linear feet of concrete capped I-wall that begins near the 

Domino Sugar Plant. There are three key points within this reach, all of them closure gates. One 

is located at the Domino Plant, one at the Port Ship Service Dock, and the last one at Mehle 

Avenue. The weighted average elevation of the top of the wall along this reach is 18.0 

(NAVD88). This reach was not damaged during Katrina. One of the closure areas and typical 

I-wall along this stretch is shown in Figure STB 21. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-24

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Figure STB 21. Concrete capped I-wall near Domino Sugar Plant (closure represents key point within this 

reach) 

Reach STB31 (Mississippi River Levee) 

This reach consists of 2,345 feet of levee that is primarily covered with paved slopes and 

roads. It begins near the Jackson Barracks and ends at the warehouse and dock area near Flood 

Street. The weighted average elevation along this stretch of levee is 19.0 (NAVD88). There are 

no key points within this reach, although there are some pipe crossings noted but not considered 

significant enough to effect the overall reliability of the reach. This reach was not damaged 

during Katrina. The end of this reach (viewed from the Reach STB30 side) is shown in 

Figure STB 22. 

Reach STB32 (Mississippi River Levee) 

This reach consists of 4,870 linear feet of levee along the Mississippi River and southeast 

side of the Inner Harbor Navigation Canal. It begins at the warehouse/dock facilities on the levee 

near Flood Street and ends at the southeast side of the Inner Harbor Navigation Canal Lock. The 

weighted average elevation of the top of this levee is 20.9 (NAVD88). There are no key points 

within this reach. This reach was not damaged during Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-25

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Figure STB 22. End Reach STB31 where levee ties into Dock Facility (Viewed from Reach STB30 side) 

Reach STB33 (Interior Local Levee) 

This reach consists of approximately 15,455 linear feet of uncapped sheetpile I-wall. The 

reach starts at the north side of the interior local levee near Pump Station #5, as shown in Figure 

STB 23. There are two key points within this reach. The first is a railroad closure near the East 

Bank Wastewater Treatment Plant, and the second is a timber closure across a road. An aerial 

view of the location where the interior local levee sheetpile wall ties into the concrete capped Iwall along the IHNC is shown in Figure STB 23. A close-up view of the start of the interior local 

levee is shown in Figure STB 24. This location is referenced in Figure STB 23. This reach had a 

weighted average top of wall elevation of 13.5 (NAVD88) prior to Katrina. Most of this reach 

was not damaged during Katrina. There was a 4,500 feet stretch of uncapped I-wall and levee 

near the parish line along this reach that was damaged and in need of repair. The general area 

where the parish line crosses the interior local levee is shown in Figure STB 1. 

Reach STB34 (Interior Local Levee) 

This reach consists of 685 feet of concrete floodwall surrounding the Jean LaFitte Pump 

Station. The weighted average top elevation of this reach is 14.0 (NAVD88) taken from a 1999 

LADOTDD physical survey of the interior local levee. There was no significant damage during 

Katrina. There are no key points within this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 23. Area showing beginning of interior local levee at the IHNC 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 24. Close-up of beginning of interior local levee 

Reach STB35 (Interior Local Levee) 

This reach consists of 5,055 linear feet of uncapped sheetpile I-wall within the levee. The 

weighted average top elevation of this reach is 13.3 (NAVD88) taken from a 1999 LADOTD 

physical survey of the interior local levee. There are a couple of pipe crossings along this reach, 

but no key points for the risk assessment. There was overtopping along this reach during Katrina, 

but no major damage. The end of the reach where the I-wall ties into an adjoining levee 

(beginning of Reach STB36) is shown in Figure STB 25. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 25. End of Reach STB35 at the beginning of Reach STB36 (interior local levee near Paris 

Road) 

Reach STB36a (Interior Local Levee) 

This reach is defined by 15,105 linear feet of levee between floodwalls. The reach begins 

approximately 2,400 feet west of Paris Road (see Figure STB25) and ends where the levee ties 

into the west end of the sheetpile I-wall leading to Pump Station #7 (Bayou Ducros Pump 

Station). There are two basic sections within this reach, the first is the short section west of Paris 

Road that had a weighted average elevation of 11.6 (NAVD88) prior to Katrina. From Paris 

Road east, the weighted average elevation of this reach was 8.4 (NAVD88) prior to Katrina. The 

main reason this reach is sub-divided was because of the work carried out by TFG was only done 

east of Paris Road with respect to raising the levee. The weighted average top of levee for this 

reach was 8.9 (NAVD88) prior to Katrina. There was significant overtopping damage to the 

section of levee east of Paris Road during Katrina. There are no key points within this reach. 

Reach STB36b (Interior Local Levee) 

This reach consists of 350 linear feet of floodwall surrounding Pump Station #7, also referred 

to at the Bayou Ducros Pump Station. The reach has 140 feet of uncapped sheetpile I-wall west 

of the discharge pipes, 70 feet of concrete T-wall around the discharge pipes, and 140 feet of 

uncapped sheetpile I-wall east of the discharge pipes. The weighted average elevation for the top 

of the 280 feet of uncapped sheetpile I-wall is 9.0 (NAVD88). The weighted average top 

elevation of the concrete T-wall is 12.5 (NAVD88). The overall weighted average top of wall 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

elevation for the entire reach is 9.7 (NAVD88). A photograph of the 3.5 feet difference in wall 

height where the sheetpile ties into the adjoining concrete T-wall at the discharge pipes is shown 

in Figure STB 26. 

Figure STB 26. T-wall and I-wall offset at Pump Station #7 along interior local levee (Note: wall offset is 

approximately 3.5 feet) 

Reach STB36c (Interior Local Levee) 

This reach consists of 20,870 linear feet of levee that begins at the east end of Pump 

Station #7 I-wall and then ends where the levee adjoins the Violet Canal. This reach of levee had 

a weighted average top elevation of 9.1 (NAVD88). There are no key points within this reach. 

This reach of levee was overtopped and damaged during Katrina. 

Reach STB37 (Interior Local Levee) 

This reach is defined by 3,888 linear feet of levee along the north side of the Violet Canal. 

The weighted average top of levee elevation along this reach was 8.1 (NAVD88) prior to 

Katrina. There are no key points within this reach. This reach was overtopped and damaged 

during Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-30

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB38 (Interior Local Levee) 

The floodwall for the Soap Factory along the Violet Canal defines this reach. It is 432 linear 

feet. The weighted average elevation of the top of the wall was 7.0 (NAVD88) prior to Katrina. 

This section of wall was overtopped during Katrina, but the wall was not heavily damaged. 

There are no key points within this reach. See Figure STB 27 for a view of this reach along the 

Violet Canal. 

Figure STB 27. Floodwall Along the North Side of Violet Canal (This is Reach STB38 for the Risk Model, 

Violet Canal is to the left) 

Reach STB39 (Interior Local Levee) 

This stretch of levee runs between the end of the concrete floodwall for the Soap Factory and 

ties into the series of buildings on top of the levee along the north side of the Violet Canal. This 

section can be seen in the background of Figure STB 27. This levee is approximately 510 feet 

long and had a weighted average top elevation of 8.5 (NAVD88) prior to Katrina. The reach was 

overtopped and damaged during Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-31

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB40 (Interior Local Levee) 

Prior to Katrina, this reach was defined by a short concrete floodwall surrounding a shrimp 

factory along the Violet Canal. This floodwall was 155 linear feet and had a weighted average 

top elevation of 7.5 (NAVD88). There are no key points within this reach. The reach was 

damaged during Katrina. 

Reach STB41 (Interior Local Levee) 

This reach is defined by 2,201 linear feet of levee along the Violet Canal, and it ends where it 

ties into Highway 46. The weighted average top elevation of this reach was 8.7 (NAVD88). 

There are no key points within this reach. This reach was overtopped and damaged during 

Katrina. 

Reach STB42 (Interior Local Levee) 

This reach starts on the south side of the Violet Canal at Highway 46 and continues to the tiein to the exterior levee along the Caernarvon Canal. While there are a few ramp and pipe 

crossings within this 55,227-ft reach; none are considered key points. The weighted average 

elevation of the top of the levee along this reach was 7.7 (NAVD88). This reach was overtopped 

and damaged during Katrina. 

STB Basin – Layout of Reaches for Risk Model by Physical 

Feature (Post-Katrina Changes by Task Force Guardian) 

Reach STB1 (INHC to Caernarvon) 

Although this section of wall did not fail during Katrina, improvements were made to this 

reach to improve its performance for stability and overtopping. A physical survey of this section 

of wall taken in November 2005 shows an average elevation of 13.0 (NAVD88) across this 

reach. The wall was designed with a free standing height of 6 feet (see Figure STB 6 with a top 

of wall elevation of 15.0 and top of levee elevation of at 9.0). TFG restored the top of levee to 

9.0 feet, thus reducing the free standing height to approximately 4.0 feet. In addition, a scour 

protection slab was placed on the protected side. This is generally 6 to 8 feet wide from the base 

of the wall at the top of the levee on the protected side. 

Reach STB2 (INHC to Caernarvon) 

This section of I-wall failed during Katrina (see Figure STB 5 for reference). The top of the 

I-wall was approximately at elevation 13.0 (NAVD88 2004.65) prior to Katrina. This wall was 

replaced by a T-wall to the authorized elevation of 15.0. In addition, scour protection was 

provided on the protected side in the form of an 8-in. concrete slab that is 8-ft wide. See 

Figure STB 28 for a photograph showing the new T-wall under construction and Figure STB 29 

for the design section that was installed. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-32

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 28. New T-wall being constructed along IHNC (Reach STB2) (looking north along the IHNC) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-33

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 29. New T-wall design section (Reach STB2) (Pre-Katrina I-wall shown in picture for reference 

with authorized elevation of 15.0 feet) 

Reach STB3 (IHNC to Caernarvon) 

The major change to this reach was that scour protection was provided on the protected side 

by means of a concrete slab that is 6 to 8 feet wide extending from the face of the wall on top of 

the levee. 

Reach STB4 (IHNC to Caernarvon) 

Improvements to this reach include reducing the free standing height of the I-wall and 

providing scour protection. The approximate weighted average for the top of this wall is 

elevation 13.3 (NAVD88). The original free standing height design varied, but generally was in 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

the 5.5 feet range with a top of levee elevation at 9.0. TFG will provide a stability berm to 

elevation 9.0 on the protected side. Therefore, the free standing height will be reduced to 

approximately 4.3 feet with the increased stability berm. In addition, a scour protection pad, as 

shown in Figure STB 30, was installed on the top of the protected side levee. 

Reach STB5 (INHC to Caernarvon) 

There were no improvements made to this reach under TFG. 

Reach STB6 (IHNC to Caernarvon) 

Scour protection was provided on the protected side of all these walls and around the tie-in to 

the levee by TFG. Figure STB 31 depicts the completed scour protection pad construction. 

Reach STB7 (IHNC to Caernarvon) 

There were no improvements made to this reach under TFG. 

Figure STB 30. New scour protection pad (typical) (photo taken at end of Reach STB4 looking back 

toward IHNC) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-35

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 31. New scour protection around Reach STB6 

Reach STB8 (IHNC to Caernarvon) 

The floodwall around Paris Road has a couple of modifications made to it by TFG. First, the 

free-standing height of the wall was returned to the as designed condition. Any location where 

the free-standing height of wall was greater than designed was reduced back to a height of 6 feet. 

Secondly, scour protection was added around this wall similar to other sections. The pad extends 

10 feet away from the vertical face of the protected side. The repairs that were made are shown 

in Figure STB 32. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-36

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 32. TFG repairs to Paris Road floodwall 

Reach STB9 (IHNC to Caernarvon) 

There were no improvements made to this reach under TFG. 

Reach STB10 (IHNC to Caernarvon) 

A scour protection pad was added around the transition areas with the adjoining levee 

sections as well as the protected side of this wall by TFG as shown in Figure STB 33. There 

were no other changes to this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-37

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 33. New scour protection around Reach STB10. (viewed from Bienvenue Control Structure 

side looking northwest)

Reach STB11 (IHNC to Caernarvon) 

This short section of levee was topped with heavy riprap between the two floodwall sections 

as part of repairs made by TFG following Katrina. This will be a significant improvement in 

scour protection for this section since there is no exposed earthen levee and the transition areas 

have been topped by a combination of concrete scour pads and heavy riprap. 

Reach STB12 (IHNC to Caernarvon) 

As shown in Figures STB11-13, this area suffered significant damage during Katrina. Two 

major repair efforts were undertaken for this set of walls as part of TFG. The first is heavy riprap 

has been placed around the northwest I-wall/T-wall sections leading to the control gated 

structure. A photograph of this repair is shown in Figure STB 34 and is also typical of the riprap 

that was placed on top of Reach STB 11 as part of the repair effort. The second major repair is a 

sheetpile diaphragm wall that has been installed to replace the failed uncapped I-wall section. 

This diaphragm wall was constructed to an elevation of 18.5 (NGVD). The basic design section 

for this sheetpile diaphragm wall is shown in Figure STB 35. Note that this sheetpile diaphragm 

repair is the same one used for the northwest walls leading to the Bayou Dupree Control 

Structure (Reach STB 20), and the section shown is actually from the Bayou Dupre drawing set. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-38

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 34. Riprap placement around Northwest Floodwall adjoining Bayou Bienvenue Control 

Structure (looking southeast alogn MRGO)

Figure STB 35. New sheetpile diaphragm wall replacing failed uncapped I-wall at Bayou Bienvenue 

Control Structure. (Note: This is same repair being done for northwest wall at Bayou Dupree)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-39

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB13 (IHNC to Caernarvon) 

This reach was rebuilt to a constructed elevation of 20.0 (NGVD) with a final design grade 

of 17.5 (NGVD). This will allow for 2.5 feet of long-term settlement. There are three locations, 

each 300 feet long, within this reach where the levee was only rebuilt to elevation 17.5 (NGVD). 

These are all at utility pipe crossings. As was noted in the pre-Katrina section, this reach failed 

and had an approximate weighted average top elevation of 17.5 (NGVD). Therefore, the reach 

will be 2.5 feet higher initially until it begins to settle over time. 

Reach STB14 (IHNC to Caernarvon) 

The uncapped I-wall that defined this section pre-Katrina was removed by TFG, and the 

entire stretch was replaced with a rebuilt levee to a constructed elevation of 20.0 (NGVD) with a 

final design grade of 17.5 (NGVD), thus, allowing for 2.5 feet of long-term settlement. There are 

two exceptions to this description within this reach. There are two separate 300-ft stretches 

where the levee was only rebuilt to elevation 17.5 (NGVD) because of utility pipe crossings. See 

Figure STB 36 for reference of these areas along this stretch. 

Figure STB 36. Profile of rebuilt levee along MRGO (Reach STB14). (Note: Jagged line represents 

elevation of levee following Katrina) 

Reach STB15 (IHNC to Caernarvon) 

This section of levee, which had an approximate weighted average top elevation of 16.4 ft 

(NGVD) prior to Katrina, was replaced with a new levee between the two control structures 

constructed to elevation 20.0 (NGVD) to allow for 2.5 feet of long-term settlement. Final design 

grade is elevation 17.5 (NGVD). Thus, the new levee section will be 3.6 feet higher than the preKatrina elevation for this levee initially and then 0.9 feet higher assuming a final design grade 

after settlement occurs. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-40

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB16 (IHNC to Caernarvon) 

The uncapped I-wall that defined this section pre-Katrina was removed by TFG, and the 

entire stretch replaced with a rebuilt levee to a constructed elevation of 20.0 (NGVD) with a 

final design grade of 17.5 (NGVD), thus, allowing for 2.5 feet of long-term settlement. 

Reach STB17 (IHNC to Caernarvon) 

This reach of levee was replaced with a new levee to a constructed elevation of 20.0 ft 

(NGVD) with a final grade elevation of 17.5 (NGVD). This will allow for 2.5 feet of long-term 

settlement. This means the new levee will be 3.5 feet higher than pre-Katrina at the time of 

construction, and its final design elevation should be approximately 1.0 foot higher. 

Reach STB18 (IHNC to Caernarvon) 

The uncapped I-wall that defined this section pre-Katrina was removed by TFG, and the 

entire stretch was replaced with a rebuilt levee to a constructed elevation of 20.0 (NGVD) with a 

final design grade of 17.5 (NGVD), thus, allowing for 2.5 feet of long-term settlement. 

Reach STB19 (IHNC to Caernarvon) 

This reach of levee was replaced with a new levee to a constructed elevation of 20.0 

(NADV88) with a final grade elevation of 17.5 (NAVD88). This will allow for 2.5 feet of longterm settlement. This means the new levee will be 1.3 feet higher than pre-Katrina at the time of 

construction, and its final design elevation will actually be lower than the pre-Katrina weighted 

average elevation. 

Reach STB20 (IHNC to Caernarvon) 

There are several changes to this structure that were made by TFG. First the northwest 

precast concrete sheetpile wall that failed during Katrina was replaced with a diaphragm 

sheetpile cell wall (similar to the one shown in Figure 35 for Bayou Bienvenue). This diaphragm 

cell wall was built to elevation 18.5 (NAVD88). In addition, the adjoining T-walls on both sides 

remained, but the surrounding earthen sections were covered with heavy riprap for scour 

protection. Finally, the southeast precast concrete sheetpile wall remained, but now has an 

earthen berm placed on both the sides of the wall and covered with heavy riprap, as shown in 

Figure 37. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-41

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 37. Repairs to Southeast adjoining walls at Bayou Dupre (viewed from southeast side looking 

northwest along the MRGO) 

Reach STB21 (IHNC to Caernarvon) 

This reach of levee was replaced with a new levee to a constructed elevation of 20.0 (NGVD) 

with a final grade elevation of 17.5 (NGVD). This allows for 2.5 feet of long-term settlement. 

This means the new levee will be 0.9 feet higher than pre-Katrina at the time of construction, and 

its final design elevation will actually be lower than the pre-Katrina weighted average elevation 

across this reach if it settles to the design elevation. 

Reach STB22 (IHNC to Caernarvon) 

The uncapped I-wall that defined this section pre-Katrina was removed by TFG, and the 

entire stretch was replaced with a rebuilt levee to a constructed elevation of 20.0 (NGVD) with a 

final design grade of 17.5 (NGVD), thus, allowing for 2.5 feet of long-term settlement. 

Therefore, the levee will be 1.0 feet higher when constructed, but will be lower if settlement 

occurs to the design grade. 

Reach STB23a (IHNC to Caernarvon) 

This section of levee was rebuilt by TFG to a constructed elevation of 20.0 (NGVD) with a 

final design grade of 17.5 (NGVD). Therefore, the “new” levee will be slightly higher (0.5 feet) 

than pre-Katrina when constructed, but lower if it settles to design grade. 

Reach STB23b (IHNC to Caernarvon) 

There were small changes made to this reach under TFG. The first 1,400 feet or so of this 

levee was topped to an elevation of 19.0 (NAVD88). The remainder only had minimal scour 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-42

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repairs where necessary, thus, overall the reach will not vary greatly from pre-Katrina 

conditions. 

Reach STB23c (IHNC to Caernarvon) 

There were no major changes planned for this section of levee under TFG. There were some 

minor scour repairs to areas that were damaged during Katrina. 

Reach STB23d (IHNC to Caernarvon) 

There were no improvements planned for this reach under TFG. 

Reach STB24 (IHNC to Caernarvon) 

There were no improvements planned for this reach under TFG. 

Reach STB25 (IHNC to Caernarvon) 

Scour protection was provided along this reach under TFG even though it was not 

overtopped or damaged during Katrina. A photograph of this protection (same general location 

as pre-Katrina condition shown in Figure STB18) is shown in Figure STB 38. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-43

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure STB 38. Scour repairs to Caernarvon sheetpile I-wall area (Note: compare to pre-Katrina condition 

shown in Figure STB 18) 

Reach STB26 (IHNC to Caernarvon) 

The only significant repair to this reach undertaken by TFG was to provide a scour protection 

pad similar to the one shown for Reach STB25 (see Figure STB 38). In addition, any spots that 

had an I-wall “stick-up” height that was greater than designed were modified to reduce free 

standing height to design levels. 

Reaches STB27 through STB33 (Mississippi River Levee) 

There were no improvements to any of these reaches planned under TFG. 

Reach STB33 (Interior Local Levee) 

Most of this 15,455-ft stretch of levee did not require any modifications by TFG. However, a 

4,500-ft section of uncapped sheetpile I-wall embedded within the levee was repaired because of 

Katrina damage. This repair essentially consisted of replacing the damaged wall and levee that 

was washed away to its pre-Katrina condition. Therefore, for the purposes of the IPET 

assessment it was assumed that no improvements to this reach were made compared to the 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-44

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

pre-Katrina condition. A photograph of the construction being carried out to repair the damaged 

section along this reach is shown in Figure STB 39. 

Figure STB 39. Repairs to interior local levee along Reach STB33 near parish border (Note: This section 

is being returned to pre-Katrina conditions) 

Reach STB34 (Interior Local Levee) 

There were no improvements to this reach planned under TFG. 

Reach STB35 (Interior Local Levee) 

There were no improvements to this reach planned under TFG. 

Reach STB36a (Interior Local Levee) 

TFG repaired Katrina damages and raised the levee east of Paris Road within this reach. 

Prior to Katrina, the weighted average top elevation east of Paris Road was 8.4 (NAVD88) based 

upon a 1999 LADOTD physical survey. TFG was granted approval on a one-time basis to raise 

this portion of the levee to elevation 10 (NAVD88 2004.65). 

Reach STB36b (Interior Local Levee) 

There were no major improvements to this reach being planned by TFG. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-45

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB36c (Interior Local Levee) 

This reach of levee was rebuilt and raised to elevation 10 (NAVD88 2004.65) under TFG. 

The pre-Katrina weighted average top of levee elevation across this reach was 9.1 (NAVD88). 

Reach STB37 (Interior Local Levee) 

This reach was rebuilt and raised to elevation 10 (NAVD88 2004.65) under TFG. The preKatrina weighted average top of levee elevation across this reach was 8.1 (NAVD88). 

Reach STB38 (Interior Local Levee) 

There were no improvements planned to this section of wall under TFG. 

Reach STB39 (Interior Local Levee) 

This 510-ft reach of levee was rebuilt by TFG and raised to elevation 10 (NAVD88 2004.65). 

Prior to Katrina, the weighted average top elevation of this reach was 8.5 (NAVD88) based upon 

a 1999 LADOTD physical survey. Figure STB 40 shows the levee being rebuilt by TFG. 

Figure STB 40. Repairs and levee raising to Reach STB39 along Violet Canal (viewed looking west from 

end of Reach STB38) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-6-46

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach STB40 (Interior Local Levee) 

This reach along the Violet Canal was repaired by TFG. The section was reinforced with 

gabion baskets in front of the building facilities, as shown in Figure STB 41. The new top 

elevation of this reach is 10 (NAVD88 2004.65) compared to elevation 7.5 (NAVD88) prior to 

Katrina. 

Figure STB 41. Gabion basket repairs to Reach STB40 (viewed from west side of reach) 

Reach STB41 (Interior Local Levee) 

This reach of levee was rebuilt and raised to elevation 10 (NAVD88 2004.65) by TFG. Prior 

to Katrina, the weighted average elevation was 8.7 (NAVD88). 

Reach STB42 (Interior Local Levee) 

This reach of levee was also rebuilt and raised to elevation 10 (NAVD88 2004.65) by TFG. 

Prior to Katrina, the weighted average elevation was 7.7 (NAVD88). 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-1

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 7 

Orleans Basin 

Field Reconnaissance and Definition of Reaches in Orleans East 

Bank Basin 

The basin for Orleans East Bank has been broken down into seven distinct sections to 

develop both reaches and features definitions for the risk model. These sections are based on 

General Design Memorandum (GDMs) published by the New Orleans District, USACE and 

updated by field reconnaissance by IPET Team 10. The Orleans East Bank Basin has been 

defined by the following sections shown in the figure below: 

1. 17th Street Canal 

2. Orleans Canal 

3. London Canal 

4. Bayou St. John Canal 

5. Lakefront Levees 

6. INHC 

7. Mississippi River Levees 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-2

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Overview of Orleans East Bank Basin 

Elevations 

All vertical elevations in this report are defined in NAVD88 2004.65 datum unless otherwise 

noted within the text. All horizontal datums are defined in State Plane Coordinates NAD83 – 

1702 Louisiana South, UTM NAD83 Zone 15, and GCS NAD83. Elevations for transitions 

features are top elevation of feature and elevation for the gates are inverts. 

Section 1: 17th Street Canal (East Side) 

Narrative 

The 17th Street Canal Floodwall cross section is typical from station 125+87.45 W/L to 

0+00.00 W/L (also Sta. 340+40.00 B/L Lakefront) with a I-wall section at elevation of 14.0 feet 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-3

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NGVD. The elevation changes from station 80+10.00 W/L to 80+40.00 W/L, where the I-Wall 

crosses under Veterans Highway and the elevation is 14.5 feet NGVD. The elevation changes 

again from station 100+86.00 W/L to 102+06.00, where the I-Wall elevation is 15.0 feet NGVD. 

There was a major failure of the I-wall floodwall protection on the east side of the 17th Street 

Outfall Canal. This failure resulted in a breach located approximately N30°01.02 and 

W90°07.28. The length of the breach is 455 feet long. Floodwall monoliths, founding levee and 

foundation materials were scoured away resulting in scour hole with an approximate bottom 

elevation -21 feet NGVD. A temporary levee was constructed up to elevation +10 NGVD to 

provide temporary flood protection. Figure 3 shows the breach along the 17th Street Canal. 

Figure 4 shows the sheetpile repair along the 17th Street Canal. 

Figure 3. Breach in the 17th Street Canal 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-4

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Figure 4. Sheetpile repair along the 17th Street Canal 

Elevation Stationing (used East Bank for OEB)

Section Start End Structure EL Section/Point Length Avg Height Weighted Average

1 0+00 7+45.59 I-Wall 14 S 745.59 14

2 7+45.59 7+52.09 T-Wall 14 P 6.5 14

3 7+52.09 8+14.09 Gate No. 3 14 P 62 14

4 8+14.09 8+20.59 T-Wall 14 P 6.5 14

5 8+20.59 80+10.00 I-Wall 14 S 7189.41 14 Reach 1 14.0

6 80+10.00 80+40.00 I-Wall Transition 14.0 - El 14.5 S 30 14.25

7 80+40.00 80+51.70 I-Wall 14.5 P 11.7 14.5

8 80+51.70 81+63.70 Veterans Hwy 14.5 S 112 14.5

9 81+63.70 100+86.00 I-Wall 14.5 S 1922.3 14.5 Reach 2 14.5

10 100+86.00 102+06.00 I-Wall Transition 14.5 - El 15.0 S 120 14.75

11 102+06.00 125+87.45 I-Wall 15 S 2381.45 15 Reach 3 15.0

End at Hammond Highway
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-5

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Elevation Feature

15 Pump Station #6 - End of 17th Street Canal 

15 Pump Station OP#10 

10.6 Gate E9 - Southern RR 

7 Gate at OP#10 Pump Station 

7 Gate north of I-10 

7.3 Gate E4 - Veterans Blvd. 

7.3 Gate E5 - Veterans Blvd 

Definition of Reaches 

Reach 1-3 (now 45) – This reach consists of I-wall on the east side of the 17th Street Canal 

at average elevation 15.5 feet with a length of approximately 12,740 feet. There are two 

transitions (Pump Stations 6 and 10) and five gates in this reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-6

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-7

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-9

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-10

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-11

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-12

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Section 2: Orleans Avenue Canal 

Narrative 

The protection on the east side of the Orleans Outfall Canal starts at the intersection of 

Lakeshore Drive and Marconi Drive. After short lengths of floodwall and cantilever sheeting, the 

protection consists of a levee that extends to Robert E. Lee Boulevard. South of this intersection, 

the protection consists of floodwall that extends to the pumping station near Interstate Highway 

610. This line of protection is in satisfactory condition and does not appear to have been 

damaged during Hurricane Katrina. 

Definition of Reaches 

Reach 4-10 (now 46-49) – These reaches consists primarily of I-walls with levees at the 

lakefront entrance on both on the east side of the Orleans Avenue Canal at average elevation 

14.0 feet with a length of approximately 9,280 feet, 3,155 feet, 9,110 feet, and 3,610 feet 

respectively. There are one transitions (Pump Stations 7) and four gates in this reach. 

Section Structure Length Elevation Section Structure Length Elevation

1 I-Wall 3404 13.6 1 I-Wall 3430 13.6

2 I-Wall 2712 13.6 2 I-Wall 2545 13.6

3 I-Wall 2627 13.4 3 I-Wall 2643 13.4

4 I-wall 180 13 Reach 4 4 I-wall 160 13 Reach 8

5 Levee 2210 13 Reach 5 5 Levee 2948 13

6 I-wall 500 12 Reach 6 6 Levee 226 12 Reach 9

7 Levee 665 13 Reach 7 7 I-wall 482 13 Reach 10
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-13

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Elevation Feature

15 Pump Station OP#7 FW - End of Orleans Canal 

7 Gate at Harrison Ave 

7 Gate at Harrison Ave 

7 Gate at Filmore Ave 

7 Gate at Filmore Ave 

Pump Station #6 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-14

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

I-610 Overpass 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-15

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Harrison Ave. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-16

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-17

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-18

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Filmore Ave. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-19

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-20

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Robert E. Lee Blvd 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-22

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-23

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Lakeshore Drive and Marconi 

Section 3: London Avenue Canal 

Narrative 

There was a major failure of the I-wall floodwall protection on the east side of the London 

Avenue Outfall Canal immediately north of Mirabeau Avenue Bridge. This failure resulted in a 

breach located approximately N30° 00.52’ and W90° 04.167. The length of the breach was 

425 feet long. Floodwall monoliths, founding levee and foundation materials were scoured away 

resulting in scour hole with an approximate bottom elevation -28 feet NGVD. A provisional 

levee was constructed up to elevation +10 NGVD to provide temporary flood protection. 

Figure 5 shows the sheetpile repair along the London Avenue Outfall Canal. 

There was a second major failure of the I-wall floodwall protection on the west side of the 

London Ave Outfall Canal immediately south of Robert E. Lee Boulevard Bridge. This failure 

resulted in a breach located approximately N30°01.218’ and W90°04.26’. The length of the 

breach was 720 feet long. Floodwall monoliths, founding levee and foundation materials were 

scoured away resulting in scour hole with an approximate bottom elevation -20 feet NGVD. A 

provisional levee was constructed up to elevation +10 NGVD to provide temporary flood 

protection. Figure 6 shows the sheet pile repair and also a failed I-wall floodwall section. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-24

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 5. Sheetpile repair along the London Ave Canal – East Side Failure 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-25

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 6. Sheetpile repair along the London Ave Canal – West Side Failure 

Definition of Reaches 

Reach 11-18 (now 50-53) – These reaches consists primarily of I-walls with levees at the 

lakefront entrance on both on the east side of the London Avenue Canal at average elevation 

14.0 feet with a length of approximately 12,130 feet, 3,880 feet, 12,765 feet and 3,030 feet 

respectively. There are two transitions (Pump Stations 3 and 4) and three gates in this reach. 

Elevation Stationing - WEST (START AT PUMP STATION #3) Elevation Stationing - EAST (START AT PUMP STATION #3)

Section Structure Length Elevation Section Structure Length Elevation

1 I-Wall 9712 13 Reach 11 1 I-Wall 9827 13.2 Reach 16

2 T-Wall 300 15 Transition PS #4 2 T-wall 240 18 Transition

3 I-Wall 1984 13.4 Reach 12 3 I-Wall 1974 13.2

4 Levee 537 12 Reach 13 Between Robert Lee and Simon 4 I-wall 672 13.1

5 I-wall 190 12.5 Between Robert Lee and Simon 5 I-wall 130 13 Reach 17

6 I-wall 92 13 Reach 14 6 Levee 3647 12.5 Reach 18

7 Levee 3276 12.5 Reach 15
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Elevation Feature

9.0 Unprotected area adjacent Pump Station #3 

14.6 Pump Station #4 West FW - Middle of London Canal 

13.6 Pump Station #3 FW - End of London Canal 

10 West CSX RR gate near Pump Station #3 

7 Gate at Filmore Ave 

7 Gate at Filmore Ave 

Pump Station #3 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pump Station #3 

I-610 overpass 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Gentilly Blvd 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-30

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-31

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-32

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Mirabeau Ave. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-33

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Mirabeau Ave. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Filmore Ave 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-35

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pump Station 4 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-36

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-37

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Robert E. Lee Blvd 

Leon C Simon Blvd 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-38

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-39

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-40

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Lakeshore Blvd 

Section 4: Bayou St. John Canal 

Narrative 

The protection on the east and west banks of the Bayou St. John Canal meets the Lakefront 

protection. This protection extends approximately 800 feet to the south where there is a closure 

structure across Bayou St. John. The closure structure has both a sector gate and slice gates. 

South of the closure structure, the protection along the banks is approximately ten feet lower. 

There are no reaches considered for this canal. The gate structure is included as both a transition 

(tie-in on both sides) and a gate structure. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-41

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Inlet to Bayou St. John 

Protection Gate Structure at Bayou St. John 

Section 5: Lakefront Levees 

Narrative 

Orleans East Bank Lakefront Levee Reaches 9 and 8, as designated in Design Memorandum 

No.13, extend from the 17th Street Canal at station 340+40 to the Orleans Avenue Canal at 

station 250+72. Starting at the east bank of the 17th Street Canal, the protection begins with a 

slide gate across Marina Boulevard. At this junction, construction work was ongoing under 

contract number W912P8-06-C-0008, titled “17th Street Canal Interim Closure Structure.” 

Contract completion date is scheduled for 1 June 2006. The flood protection along Marina 

Boulevard consists of an I-wall with a series of street gates that provide access through the line 

of protection. The I-wall transitions into the levee just beyond the intersection at Lake Shore 

drive at longitude-latitude coordinates 30o1’19” North by 90o06’47” West. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-42

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The typical cross section for the Lakefront Levee has a ten-foot top width with vegetated side 

slopes. The crest has an undulant profile, and there appears to be a low spot in the protection 

where Canal Boulevard intersects the levee. Design Memorandum No. 13 indicates the net grade 

for the levee crest to be elevation 18 feet NGVD. At the Orleans Outfall Canal, the Lakefront 

Levee drops approximately four feet and transitions into the west levee for the Orleans Outfall 

Canal. At this junction, construction work was ongoing under Contract Number W912P8-06-C –

0097, titled “Interim Closure Structure, Orleans Avenue.” A review of the contract documents 

indicates the at the west levee will be raised from Lake Shore Drive to the new Closure Structure 

and the southern limit of the work will extend to West Robert E. Lee Boulevard. Contract 

completion date is scheduled for 1 June 2006. South of Robert E. Lee Boulevard, the protection 

on the west side of the Orleans Outfall Canal consists of floodwall up to the pumping station at 

29o59’40” North by 90o06’02” West. South of the pumping station, the canal is contained in an 

underground conduit. 

Orleans East Bank Lakefront Levee Reach 7 extends from the Orleans Avenue Canal at 

station 246+37 to the Bayou St. John Canal at station 199+42. On the east side of the Orleans 

Outfall Canal, there is a street gate marked “L8” across Marconi Drive and then an approximate 

20-foot length of sheeting that transitions into the typical levee cross-section that Design 

Memorandum No. 13 indicates to have a net crest elevation of 17.5 feet NGVD. The Lakeshore 

Levee then continues along the south side of Lake Shore Drive until coordinate point 30o1’39” 

North by 90o05’17” West where the alignment shifts to the north side of Lake Shore Drive where 

a flood wall serves as the protection. The point where the protection intersects Lake Shore Drive 

appears to be a low point. The floodwall transitions back to the typical levee cross section at 

coordinate point 30o1’39” North by 90o05’-15” West where minor erosion was observed on the 

lakeside slope. The typical levee cross section then continues toward Bayou St. John where it 

transitions back to a floodwall that ties into the bridge abutments where Lake Shore Drive passes 

over Bayou St. John. 

Orleans East Bank Lakefront Levee Reach 6 extends from the Bayou St. John Canal at 

station 196+50 to the London Ave Canal at station 163+98. Starting at the east bridge abutment, 

the protection consists of floodwall up to Lake Terrace Drive where it then transitions back to 

the typical levee cross-section that Design Memorandum No. 13 indicates was to be constructed 

to a net elevation of 18 feet NGVD. The transition area appears to be a low point in the 

protection. The levee section runs along the north side of Lake Shore Drive until it approaches 

the London Avenue Outfall Canal where it then crosses to the south side of Lake Shore Drive 

where it ties into the London Avenue west levee. Corps survey markers were found in this area 

with spot elevations of 14.47 feet NGVD on the crest of Lakeshore Levee and elevation 

12.84 feet NGVD on the crest of the west levee along the canal. The point where the protection 

crosses Lake Shore Drive appears to be a low point. 

Orleans East Bank Lakefront Levee Reach 4 extends from the London Ave Canal at station 

161+00.18 B/L to the Inner Harbor Navigational Canal at station 0+00.00. This reach is typically 

levee. The section from station 136+13.19 B/L to station 102+23.16 B/L has been removed and a 

new floodwall has been constructed on the back side of the University of New Orleans. The 

original DM showed the floodwall on the front side of what is not the University of New 

Orleans. This area includes floodwall and levee. Figure 7 shows the new floodwall and levee. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-43

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The majority of the damage in this reach is lake side erosion, scour at the base of the floodwalls, 

and damaged floodwalls. There are also trees and bushes growing adjacent to and on the levee. 

Figure 8 shows the bushes on the levee. 

Figure 7. Floodwall and Levee behind the University of New Orleans 

Reach 23-32 (now 54-60) – These reaches consist primarily of levees on the lakefront with 

walls at certain protection areas. Elevations range from 12 to 18 feet depending upon the section. 

The lengths of the reaches are shown with their respective elevations. There are eleven 

transitions and eighteen gates in this reach. 

Reach Length Elevations

54 2,925 12.00 

55 6,310 18.00 

56 9,940 17.00 

57 2,380 16.50 

58 3,220 16.50 

59 7,605 16.50 

60 1,155 14.40 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-44

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Elevation Stationing

Section Start End Structure EL Section/Point Length Avg Height W eighted Average

1 0+00 1+13.57 I-W all El 14.5 S 113.57 14.5

2 1+13.57 1+21.07 I-W all El 14.0 P 7.5 14

3 1+21.07 1+51.07 Ramp formerly Gate No 1 El 10.0 P 30 14

4 1+51.07 1+58.57 I-W all El 14.0 P 7.5 14

5 1+58.57 2+54.51 I-W all El 14.5 S 95.94 14.5 Reach 32 14.41

6 2+54.51 5+60.00 Levee El 14.0 S 305.49 14

7 5+60.00 5+74.37 Levee Transition El 14.0 - El 15.0 S 14.37 14.5

8 5+74.37 7+04.00 Levee El 15.0 S 129.63 15 Reach 31 14.30

9 7+04.00 7+62.87 I-W all El 15.5 S 58.87 15.5

10 7+62.87 7+69.37 I-W all El 15.0 S 6.5 15

11 7+69.37 8+02.37 Gate No. 2 El 8.0 S 33 15

12 8+02.37 8+08.87 I-W all El 15.0 S 6.5 15

13 8+08.87 8+59.15 I-W all El 15.5 S 50.28 15.5

14 8+59.15 8+63.15 I-W all El 15.0 S 4 15

15 8+63.15 8+91.15 Gate No. 3 El 10.75 S 28 15

16 8+91.15 8+95.15 I-W all El 15.0 S 4 15

17 8+95.15 9+88.10 Seabrook Bridge El 19.19 S 92.95 19.19

18 9+88.10 14+31.42 I-W all El 15.5 S 443.32 15.5

19 14+31.42 14+40.23 Levee El 15.0 S 8.88 15 Reach 30 15.90

20 14+40.23 15+30.23 Levee Transition El 15.0 - El 18.0 S 90 16.5

21 15+30.23 15+85.23 Levee El 18.0 S 55 18

22 15+85.23 15+90.23 I-W all El 18.5 S 5 18.5

23 15+90.23 16+08.23 I-W all Transition El 18.5 - El 18.0 S 18 18.25

24 16+08.23 16+27.23 I-W all El 18.0 S 19 18

25 16+27.23 16+36.23 I-W all El 18.5 S 9 18.5

26 16+36.23 16+70.23 Gate No. 4 El 12.25 S 34 17.5

27 16+70.23 16+79.18 I-W all El 18.5 S 8.95 18.5

28 16+79.18 16+95.46 I-W all El 18.0 S 16.28 18

29 16+95.46 17+13.46 I-W all Transition El 18.0 - El 18.5 S 18 18.25

30 17+13.46 17+18.46 I-W all El 18.5 S 5 18.5

31 17+18.46 20+00.00 Levee El 18.0 S 281.4 18

32 20+00.00 20+05.00 I-W all El 18.5 S 5 18.5

33 20+05.00 20+22.25 I-W all Transition El 18.5 - El 18.0 S 17.25 18.25

34 20+22.25 20+33.43 I-W all El 18.0 S 11.18 18

35 20+33.43 20+40.43 I-W all El 17.5 S 7 17.5

36 20+40.43 20+62.43 Gate No. 5 El 12.5 S 22 17.5

37 20+62.43 20+69.43 I-W all El 17.5 S 7 17.5

38 20+69.43 20+76.30 I-W all El 18.0 S 6.87 18

39 20+76.30 20+93.55 I-W all Transition El 18.0 - El 18.5 S 17.25 18.25

40 20+93.55 20+98.55 I-W all El 18.5 S 5 18.5

41 20+98.55 42+10.0 Levee El 18.0 S 2111.45 18

42 42+10.0 42+16.00 Ramp No 1 - Levee Transition El 18.0 - El 14.5 P 6 16.25

43 42+16.00 42+26.00 Ramp No 1 - Levee El 14.5 P 10 14.5

44 42+26.00 42+62.00 Ramp No 1 - Roadway El 14.5 P 46 14.5

45 42+62.00 42+68.00 Ramp No 1 - Levee El 14.5 P 6 14.5

46 42+68.00 42+80.00 Ramp No 1 - Levee Transition El 14.5 - El 18 P 12 16.25

47 42+80.00 78+55.24 Levee El 18.0 S 3575.24 18 Reach 29 17.94

48 78+55.24 78+65.24 Ramp No 2 - Levee Transition El 18.0 - El 14.5 P 10 16.25

49 78+65.24 78+70.24 Ramp No 2 - Levee El 14.5 P 5 14.5

50 78+70.24 78+98.24 Ramp No 2 - Roadway El 14.5 P 28 14.5

51 78+98.24 79+36.24 Ramp No 2 - Shoulder El 15.0 P 38 15

52 79+36.24 79+64.24 Ramp No 2 - Roadway El 14.5 P 928 14.5

53 79+64.24 79+75.24 Ramp No 2 - Levee El 14.5 P 11 14.5

54 79+75.24 79+81.24 Ramp No 2 - Concrete Capping El 18.0 P 6 18

55 79+81.24 91+50.00 Levee (Now Floodwall) El 19.5 S 1168.76 19.5 Reach 28 17.19

56 91+50.00 94+60.00 Levee Transition El 19.5 - El 17.0 S 310 18.25

57 94+60.00 102+23.16 Levee (Ramp 3 included) El 17.0 S 763.16 17 Reach 27 17.36

58 102+23.16 136+13.19 FLOODW ALL AT UNO - 1984 3390.07 20 Reach 26 20.00

59 136+13.19 159+70.00 Levee El 17.5 S 2356.81 17.5 Reach 25 17.50

60 159+70.00 163+98.15 London Ave. Outfall Canal P - - - -

61 163+98.15 166+38.00 Levee El 18.0 S 239.85 18

62 166+38.00 167+02.00 Ramp No. 5 (SEE PLATE 24 FOR DETAILS) P 64 15.5

63 167+02.00 196+50.00 Levee El 18.0 S 2948 18 Reach 24 17.95

64 196+50.00 199+41.52 Bayou St. John P - - - -

65 199+41.52 203+18.00 Levee El 17.5 S 376.48 17.5

66 203+18.00 204+28.30 Gate No 8 (SEE PLATE 17 FOR DETAIL) P 110.3 17.5

67 204+28.30 218+14.5 Levee (with 300 ft concrete capped wall at Ramp 6) El 17.5 S 1386.2 17.5

68 218+14.5 218+77.5 Ramp No. 6 (SEE PLATE 24 FOR DETAILS) P 63 14.5

69 218+77.5 244+59.81 Levee El 17.5 S 2582.31 17.5

70 244+59.81 246+10.04 Gate No 9 Marconi (SEE PLATE 18 FOR DETAIL) P 150.23 17.5

71 246+10.04 246+37.18 Levee (now concrete capped floodwall) El 17.5 S 27.14 17.5 Reach 23 17.46
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-45

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Elevation Feature

13.8 Gate at W. Roadway Street 

14 WLT - OM 10/11 

7.55 Gate 10 - Topaz Dr 

13.8 NB Ramp at Canal Blvd 

13.8 SB Ramp at Canal Blvd 

12.05 Gate 9 - Marconi Dr. 

16 Ramp 6 - Lakeshore Dr. 

16 Ramp Lake Terrace Dr. 

16 Ramp 5 - Lakeshore Dr. 

16.8 Ramp 4 - Lakeshore Dr. 

16.8 Ramp 3 - Lakeshore Dr. 

16.3 Ramp 2 - Franklin Ave - double wide ramp 

14.5 Ramp 1 - Leroy Johnson Drive 

11.8 Gate 5 - Navy Reserve 

11.55 Gate 4 - Navy Reserve 

14 WLT O 16/15 

10.05 Road Gate 3 

7.3 RR - Gate 2 

7 Gate at Filmore Ave 

7 Gate at Filmore Ave 

7 Gate at Leon C Simon Blvd 

8.8 W. Roadway St 

8.8 Gate 15 - Into Marina Parking 

8.8 Gate 14 - Into Marina Parking 

8.8 Gate 13 - Into Marina Parking 

8.8 Gate 12 - Entrance to Marina 

6.05 Gate 11 - Lakeshore Dr. 

7.55 Gate 10 - Topaz Dr 

12.05 Gate 9 - Marconi Dr. 

-5 Bayou St John Floodgate 

10 Gate 3 UNO 

10 Gate 2 UNO 

10 Gate 1 UNO 

11.8 Gate 5 - Navy Reserve 

11.55 Gate 4 - Navy Reserve 

10.05 Road Gate 3 

7.3 RR - Gate 2 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-46

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 32 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-47

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 31 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-48

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 30 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-49

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 29 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-50

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 28 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-51

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 27 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-52

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 26 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-53

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 25 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-54

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 25 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-55

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 25 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-56

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 25 

Reach 26 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-57

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 27 

Reach 27 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-58

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 28 

Reach 29 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-59

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 29 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-60

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 30 and 31 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-61

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 31 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-62

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 31 and 32 

Section 6: Inner Harbor Navigation Canal (INHC) 

Narrative 

The Inner Harbor Navigational Canal (IHNC) reach extends along the west side of the IHNC 

from the lock at St. Claude Avenue northward to Lake Pontchartrain. Flood protection along this 

reach consists of levee and floodwall. From Station 0+00.00 to Highway 90 (Station 118+85.00) 

the floodwall sustained minor damage in the form of scour along its base. Figure 8 shows the 

repair to the floodwall scour. The Levee and floodwall between Hwy 90 (Station 118+85.00) and 

Florida Ave (Station XX+XX.XX) sustained scour damage. In this area, the floodwall breached 

at N29 59.315 and W90 01.612. Figure 9 shows the repairs in the breached area. The floodwall 

between Florida Avenue and the Lock sustained damage along the base of the floodwall in the 

form of scour. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-63

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Repair to Scour Along Floodwall. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-64

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Repairs Along the IHNC Where The Floodwall Breached Levee Section. 

Definition of Reaches 

Reach 33-41 (now 61-65)– These reaches consists primarily of I-walls along the canal with a 

levee section at the Port of New Olreans facility. The elevations of the I-walls and lengths of 

reaches are shown below. There are fourteen transitions (Pump Stations 3 and 4) and thirty five 

gates in this reach. 

Reaches Length Elevations

61 9,095 13.50 

62 9,170 13.80 

63 1,490 13.80 

64 8,390 13.80 

65 875 20.10 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-65

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Elevation Feature

13.25 Ramp east of France Road near Hickey Bridge 

14.75 France Road Ramp near Chef Mentuer Hwy 

6.25 Road Gate 7W 

7.4 RR Gate 8W 

9 Ramp to Bridge 

4.75 RR Gate 9W 

14.75 France Rd Ramp 

4.25 Gate 10W - RR open 

Elevation Stationing

Section Start End Structure EL Section/Point Length Avg Height Weighted Average

1 31+05 46+00 I-Wall 14 S 1495 14 Reach 33 14.00

2 46+00 46+30 I-Wall Transition 14.0 - 14.25 30 14.125

3 46+30 61+00 I-Wall 14.25 1470 14 Reach 34 14.00

4 61+00 61+30 I-Wall Transition 14.25 - 14.5 30 14.375

5 61+30 76+00 I-Wall 14.5 1470 14.5 Reach 35 14.50

6 76+00 76+30 I-Wall Transition 14.5 - 14.75 30 14.625

7 76+30 90+70 I-Wall 14.75 1440 14.75 Reach 36 14.75

8 90+70 91+00 I-Wall Transition 14.75 - 15.0 30 14.875

9 91+00 106+25 I-Wall 15 1525 15

10 106+25 106+57 France Road Ramp 14.5 32 14.5

11 106+57 106+84.5 I-Wall 15 27.5 15

12 106+84.5 109+81.5 I-Wall 14.5 297 14.5

13 109+81.5 110+37.5 Gate 1W / T-Wall 14 56 14

14 110+37.5 112+15 I-Wall 14.5 177.5 14.5

15 112+15 112+56 Gate 2W / T-Wall 14 41 14

16 112+56 116+53 I-Wall 14.5 397 14.5

17 116+53 118+85 I-Wall 15 232 15

18 118+85 119+59 Hwy 90 17 74 17

19 119+59 121+78.5 I-Wall 14.5 219.5 14.5

20 121+78.5 122+07.5 Gate 3W 14 29 14

21 122+07.5 124+88.5 I-Wall 14.5 281 14.5

22 124+88.5 125+17.5 Gate 4W 14 29 14

23 125+17.5 128+41.5 I-Wall 14.5 324 14.5

24 128+41.5 128+70.5 Gate 5W 14 29 14

25 128+70.5 130+53.5 I-Wall 14.5 183 14.5

26 130+53.5 130+82.5 Gate 6W 14 29 14

27 130+82.5 132+00 I-Wall 14.5 117.5 14.5

28 132+00 135+31 I-Wall 15 331 15

29 135+31 136+10 Gate 7W 14 79 14

30 136+10 136+27 T-Wall 14 17 14

31 136+27 136+94.5 Gate 8W 14 67.5 14

32 136+94.5 137+42 I-Wall 15 46.5 15

33 137+42 137+72 I-Wall 14.5 30 14.5

34 137+72 141+20 Levee 15 348 15

35 141+20 143+94 I-Wall 15 274 15

36 143+94 144+01 T-Wall 14 7 14

37 144+01 144+48 Gate 9W 14 47 14

38 144+48 145+39 I-Wall 15 91 15

39 145+39 145+76 I-Wall 14.5 37 14.5

40 145+76 148+28 Levee 15 252 15

41 148+28 210+10 I-Wall 15 6182 15

Double Hung Gates (France Rd Parkw 80

42 210+10 210+89 T-Wall 14 79 14

43 210+89 211+03 Gate 10W 14 14 14

44 211+03 211+17 T-Wall 14 14 14

45 211+17 211+46 I-Wall 15 29 15

46 211+46 211+81 I-Wall 14.5 35 14.5 Reach 37 14.88

France Road Ramp 14.5

47 211+81 226+44 Levee 15 1463 15 Reach 38 15.00

48 226+44 226+60 I-Wall 15 16 15

49 226+60 235+77 T-wall 14 917 14 Reach 39 14.02

Start at Pump Station #14

50 390 I-wall/Twall 15 Transition Pumping Station #14

51 2293 I-wall 13 Reach 40

53 875 Levee 20 Reach 41 IHNC Lock
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-66

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

14.75 France Road Ramp 

2.25 Gate in Levee at Port of NO 

2.25 Gate in Levee at Port of NO 

6 Pump Station near Florida Ave Bridge 

14 LWT OM 22/23 

5 Gate in France Road 

7.25 Gate 1W road closed 

7.55 Gate 2W RR open 

9.5 Gate 3W access open 

11.25 Gate 4W access open 

11.25 Gate 5W access open 

11.25 Gate 6W access open 

6.25 Gate 7W road closed – damaged 

7.4 Gate 8W RR open – Failed during Katrina 

4.75 Gate 9W RR open – damaged 

4.25 Double Gates - France Rd Parkway – damaged 

4.25 Gate 10W - RR open – damaged 

4.25 Gate in levee at Port of NO – Failed during Katrina 

4.25 Gate in levee at Port of NO – Failed during Katrina 

2.5 Gate in pier access – no damage 

2.5 Gate in pier access – no damage 

6 Gate next to pump station 

2.25 Road closed – Florida Avenue Bridge Gate W20 

2 RR closed 

2.25 RR open – Gate W21 

2.75 Road open – Florida Ave Bridge 

7.45 Road open – Florida Avenue Wharf 

4.45 RR open – Florida Avenue Wharf 

7.25 Road open – Florida Avenue Wharf 

8 Road open 

8.25 Road open 

7.5 Road open 

1.75 Road open – Gate W6 

2.35 RR closed – Gate W5 

5.35 Road open – Gate W4 

5.35 RR Access closed – Gate W3 

3.5 Road open – Gate W2 

2.25 Road open – Gate W1 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-67

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 33 

Reach 33 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-68

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 33 

Reach 33 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-69

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 33 

Reach 33 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-70

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 33 

Reach 34 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-71

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 34 

Reach 35 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-72

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 35 

Reach 35 and 36 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-73

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 36 

Reach 36 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-74

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 36 

Reach 36 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-75

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 36 

Reach 36 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-76

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 36 

Reach 36 and 37 

Reach 37 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-77

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 37 

Reach 38 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-78

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 39 

Reach 39 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-79

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 39 

Reach 39 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-80

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 40 

Reach 40 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-81

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 40 

Reach 40 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-82

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 40 

Reach 40 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-83

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 40 

Reach 41 

Section 7: Mississippi River and Levees (MRT) 

References 

General design information was not available for the MRT walls and levees within Orleans 

Parish at the time of the report. Information was gathered from discussion with the Orleans 

Parish Levee Board and from the field recon. 

Narrative 

The MRT connects to the IHNC at the navigation lock facility. The MRT structures consist 

of levees into the INHC lock, a series of flood walls (as shown below) that surround the City of 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-84

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

New Orleans up to the Port of New Orleans and a MRT levees from the Port of New Orleans to 

the Jefferson/Orleans Parish border. 

Typical Area on the Mississippi River East Levee and Floodwall. 

Definition of Reaches 

Reach 42-50 (now 66-71) – These reaches consists primarily of I-walls with levees at the 

lakefront entrance on both on the east side of the London Avenue Canal at average elevation 

14.0 feet with a length of approximately 12,130 feet, 3,880 feet, 12,765 feet and 3,030 feet 

respectively. There are two transitions (Pump Stations 3 and 4) and three gates in this reach. 

Reaches Length Elevations

66 1,980 21.50 

67 8,915 22.50 

68 25,450 23.60 

69 10,780 24.30 

70 14,180 24.80 

71 3,350 25.80 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-85

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Elevation Feature

7.5 Road Access — Off Poland Ave - Navy Complex 

7.5 Road Access — Off Poland Ave - Navy Complex 

7.5 RR Gate — Pauline St Wharf 

7.5 Road Access — Pauline St Wharf 

7.5 RR Gate — Pauline St Wharf 

7.5 RR Gate — Off Charles St. – photo 

7.5 Road Access — Off Charles St. 

7.5 Road Access — Off Charles St. 

7.5 RR Gate — Press St. Wharf 

7.5 Road Access — Esplande St and Wharf 

7.5 Road Access — To riverfront parking off N Peters 

7.5 Road Access — To riverfront parking - St Peters 

7.5 Road Access — To riverfront parking – Toulouse 

7.5 Road Access — To riverfront parking - St. Louis 

7.5 Road Access — To riverfront parking - Conti St 

7.5 Road Access — To riverfron parking – Bienville 

7.5 Pedestrian Crossing — North end of Riverwalk 

7.5 Road Access — Convention Center openings 

5 Road Access — Henderson Street 

5 Road Access — Race Street 

5 Road Access — Orange Street 

5 Road Access — Celeste St – photo 

5 Road Access — Port of NO - near Felicity St 

5 Road Access — Port of NO - 3rd St 

5 Road Access — Port of NO - Washington St 

5 RR Gate — Port of NO - across from 9th ST 

5 Road Access — Port of NO 

5 RR Gate — Port of NO 

7.5 Road Access — Port of NO - Louisana Ave 

7.5 Road Access — Port of NO 

7.5 RR Gate — Port of NO - Napoleon Ave 

7.5 Road Access — Port of NO - Wharehouse Rd 

7.5 Road Access — Port of NO 

7.5 Road Access — Port of NO - Coffee Dr 

Section Structure EL Length Section/Point W eighted Elevations

1 Levee 21 1774 S Reach 42 21

2 I-W all 17 19661 S Reach 43 17

3 I-W all 20 2951 S Reach 44 20

4 I-wall 17 1469 S

5 I-W all 14 8371 S

6 I-W all 17 3437 S Reach 45 15.1

7 I-W all 20 1843 S

8 I-W all 18 5920 S

9 I-W all 20 1902 S Reach 46 18.8

10 I-W all 24 2961 S Reach 47

11 Levee 24 5917 S Reach 48 24

12 Levee 20 2732 S Reach 49 20

13 Levee 24 4526 S Reach 50 24
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-86

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

7.5 Road Access — Port of NO - Leake Ave 

7.5 Road Access — Port of NO - Henry Clay Dr 

Reach 42 

Reach 42 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-87

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 43 

Reach 43 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-88

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 43 

Reach 43 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-89

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 43 

Reach 43 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-90

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 43 

Reach 43 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-91

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 43 

Reach 43 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-92

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 44 

Reach 44 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-93

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 44 

Reach 44 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-94

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 45 

Reach 45 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-95

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 45 

Reach 45 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-96

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 45 

Reach 45 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-97

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 45 

Reach 45 

Reach 45 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-98

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 45 

Reach 45 

Reach 46 and 47 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-99

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 47 

Reach 47 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-100

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 47 

Reach 47 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-101

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 47 and 48 

Reach 48 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-102

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 48 

Reach 48 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-103

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 48 

Reach 48 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-104

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 48 

Reach 48 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-105

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 48 

Reach 48 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-106

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 49 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-107

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reach 50 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-108

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Orleans West Bank 

The Orleans Parrish West Bank Basin is composed of two sub-basins (Figure 1). These are 

located on either side of the upper end of the Algiers. OW1, on the west side of Algiers Canal, 

was designed as part of the Algiers Canal to Hero Canal Project. OW2, on the east side of the 

Algiers Canal, was designed as part of the Harvey Canal to Algiers Project. 

Figure 1. Orleans Parish West Bank Basin with sub-basins (OW1 and OW2) 

OW1 

Orleans West Bank – OW1 (Algiers Canal to Hero Canal Project) 

Sub-basin OW1, as shown above, is located on the west bank of the Mississippi River in 

Orleans Parish and is generally bounded by a portion of the Algiers Canal, the Mississippi River 

and the Orleans-Plaquemines Parish line. Topography is flat with ground elevations ranging 

from +5 feet NGVD on the alluvial ridges along the Mississippi River to –7 feet NGVD in the 

interior of the area. Approximately 40 percent of the area is below sea level. The surface area is 

4.7 square miles. The sub-basin area is protected by 15.0 miles of levees and floodwalls. There 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-109

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

are no floodgates, drainage structures or control structures in the protection system. There is one 

pumping station that drains the protected area (NOS&WB Pumping Station #11 at location 

29.90962 -89.978). 

Segment 1 is a low, all earth (clay) non-federal levee separating Orleans Parish from 

Jefferson Parish. It extends from the main line Mississippi River levee (MRL) inside the US 

Coast Guard Station to the Algiers Canal levee with an elevation of 3–4 ft NVGD. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-110

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

MRL and back levee intersection 

inside the US Coast Guard Station 

Back levee begins 

US Coast Guard Station 

Mississippi River Levee 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-111

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Back levee continues inside 

US Coast Guard Station 

Back levee crossed by roadway inside 

US Coast Guard Station. Roadway is 2 feet 

lover than levee. No closure structure 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-112

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Back levee continues outside US Coast Guard toward GIWW Station. Crown is rutted 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-113

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Drainage structure under Back Levee outside US Coast Guard station. Screw gate closure on culvert 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-114

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Back Levee intersects Hwy 406 looking toward the 

GIWW and General De Gaulle Bride overpass. 

Back Levee intersects Hwy 406. 

No closure at Hwy 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-115

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Segment 2 is the East bank of the Algiers Canal levee that extends between the local interior 

levee (Segment 1) and the Algiers Lock. This 9.5 ft NVGD clay levee is interrupted by a 

floodwall segment that crosses in front of NOS&WB Pumping Station #11. 

Back Levee passed under General De Gaulle onramp from 

Hwy 406. Buckling of concrete slab on levee under ramp. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-116

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Abandoned pipe crossing through the Algiers Canal 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-117

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Typical Algiers Canal levee; no armor 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-118

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

South end of Algiers Lock 

Segment 3 is the Orleans West Levee District Mississippi River levee. This levee segment 

closes the North and East side of the sub-basin. It extends from the GIWW and Mississippi River 

intersection to where it intersects the interior levee (Segment 1) inside the US Coast Guard 

Station. The MRL is an all clay levee with crushed stone surfacing on the 10-ft wide crown at 

elevation 22 ft NVGD. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-119

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Mississippi River and GIWW intersection 

GIWW 

Mississippi River 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-120

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Pipe crossing along MRL just east of GIWW and MRL intersection 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-121

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Barge sitting on foot of MRL 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-7-122

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

OW2 

Orleans West Bank – OW2 (Harvey Canal to Algiers Canal) 

Sub-basin OW2, as shown in figure above, is located on the west bank of the Mississippi 

River in Orleans Parish. It is generally bounded by the Mississippi River, the Algiers Canal, and 

the Orleans-Jefferson-Plaquemines Parish boundary. The topography is flat with ground 

elevations ranging from +10 feet NGVD on the alluvial ridges along the Mississippi River to 

-5 feet NGVD in the interior of the area. Approximately 25 percent of the area is below sea level. 

The surface area is 6.3 square miles and the population is approximately 57,000. The sub-basin 

is protected by 12.6 miles of levees and floodwalls. 

Segment 1 is a clay, non-federal levee that begins at the Greater New Orleans Bridge 

(US 90) crossing of the Mississippi River Levee and runs along the Orleans-Jefferson Parish line 

to the Algiers Canal levee, near the NOS&WB Pump Station #13. This interior levee is 

approximately 4 miles long and is at elevation 3–4 ft NGCD. 

Typical MRL. 

10 ft crown of grass and stone 

Concrete armor on 

River side of MRL 
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Segment 2 is the West bank of the Algiers Canal levee (GIWW) that extends between the 

local interior levee and the Algiers Lock. This clay levee is interrupted by a floodwall segment 

that crosses in front of NOS & WB Pumping Station #13. It is 1.8 miles long at elevation 9.5 ft 

NGVD. 

Segment 3 is the Orleans West Levee District Mississippi River Levee extending from the 

Algiers Canal Lock west to the Orleans Parish Line (beneath the Greater New Orleans Bridge, 

US 90), completing the sub-basin. This MRL is a predominately all clay levee with small 

reaches of short concrete I-Walls atop the clay levee base. It is 6.8 miles long at elevation 

22-23.5 ft NGVD. There are no floodgates, drainage culverts or control structures in the 

protection system. There is one pumping station that drains the protected area (NOS & WB 

Pumping Station #13 at location 29.8959, -89.9978). 

Risk Model Idealization 

The West BankHPS was descretized into two sub-basins (OW 1 and OW 2) as shown in 

Figure 2. The sub-basins were defined to correspond to the known interior drainage areas. This 

reach idealization follows from the basin description information presented above, which was 

collected from project documents and field inspections. Figure 3 shows the elevations for the 

Orleans Parish West Bank HPS: Pre-Katrina—at the time of Katrina, Current—as of 

1 June 2007, and the Authorized system (authorized at the time of Katrina). 
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Figure 2. Orleans Parrish West bank Basin reaches (HA1,HA 5-HA9) and sub-basins (OW 1 and OW 2) 

definition for use in the risk model. 
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Figure 3. Elevations for the Orleans Parrish West Bank (OW1 and OW2) for the Pre-Katrina HPS (in 

place when Katrina occurred), the Current HPS (as of 1 June 2007), and the Authorized HPS 

(authorized at the time Katrina occurred) (continued) 
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Figure 3. Continued 
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Figure 3. Concluded 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-1

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Appendix 8 

Hazard Analysis 

Introduction 

This appendix includes several components. First is an overall discussion of the Hazard 

Analysis modeling effort, describing the surge (ADCIRC) and wave (WAM, STWAVE) 

modeling. That first section is followed by a subsidiary section labeled Appendix 8-1 that 

defines the many acronyms used in the modeling discussion. This is followed by a major section 

labeled Appendix 8-2 consisting of the Whitepaper of Donald T. Resio, ERDC-CHL, which 

forms the basis of the storm statistics and JPM methods discussed in the main body of the report. 

Owing to its prominence in the discussion, it has, for convenience, been referred to as R2007. 

Note that R2007, itself, includes several appendices; these are identified as Appendices A-G 

within Appendix 8-2. Finally, a discussion of the rainfall model is included here as 

Appendix 8-3. 

Hazard Analysis 

The hazard analysis required for the risk assessment was based upon the hurricane modeling 

conducted by a team of Corps of Engineers, FEMA, NOAA, private sector and academic 

researchers working toward the definition of a new system for estimating hurricane surges and 

waves. Following is a discussion of the processes used by this team and the steps taken by the 

risk team to incorporate the results into the risk analysis. 

The hurricane hazard definition required as input to the risk analysis involved several steps: 

1. Selection of the methodology to be used for estimating surges and waves 

2. Determination of hurricane probabilities 

3. Production of the ADCIRC grid models for the different HPS configurations. 

4. Production of the computer system for development of the large number of hydrographs 

required by the risk model. 

5. Formatting of the ADCIRC/STWAVE hydrographs for input to the risk computer model 

6. Determination of the rainfall volumes expected for each hurricane 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-2

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The hurricane modeling resulted in a total of 152 storms that were used in the risk analysis. 

Of these storms, frequencies were developed by the storm team for only 77 storms. The storm 

parameters and there frequency of occurance are shown in Table 8-1. 

Methodology for Hurricane Modeling 

At least five methods have been applied in past studies of environmental extremes due to 

hurricanes in the United States: 

1. Formulation of design storm events 

2. Estimates based only on historical storms 

3. The empirical simulation technique (EST) 

4. The joint probability method (JPM) and 

5. The Empirical Track Model 

All of the methods referenced above have different strengths and weaknesses for various 

applications. Appendix 2 (R2007) includes a complete discussion of each of these methods. It 

considers includes the following topics: 

1. Potential extensions of probabilistic methods to future hurricane surges, 

2. The JPM with Optimal Sampling (JPM-OS); and 

3. A computational methodology for effective simulation of storm surges for hurricane 

inundation studies. 

In each method, it is important to understand that two different statistical measures are 

required to characterize the expected extremes over an interval of time. The first of these is the 

measure of the expected values of the distribution and the second is a measure of expected 

dispersion around these central values. 

Hydrograph Production System 

Overview 

On-going projects for storm-surge and inundation mapping along the US Gulf of Mexico 

coast require the computation of many simulations using state-of-the-art numerical models and 

high-performance computational facilities. In this context, the storm-surge model ADCIRC and 

spectral wave model STWAVE are used in a “coupled” scenario to improve upon the storm 

surge solutions provided by ADCIRC. Coupling of the models is done through input/output file 

exchange between the numerical models. The water level and wind field from ADCIRC are 

provided in the appropriate form for the spectral wave model STWAVE. STWAVE computes 

the high-frequency wave field, including the wave radiation stress associated with the storm 

(wind) and simulated water level. This wave radiation stress is then passed to the next ADCIRC 

simulation as an additional forcing term, in conjunction with the storm wind and pressure field. 

This model communication through file I/O achieves a first-order coupling of the non-linear 

interaction between the wind-induced storm-surge and the wind-driven higher frequency wave 

fields. 
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The primary difficulty with the large number of compute-resource-intensive simulations 

required by these projects is ensuring that each simulation uses the correct files and that the 

compute-jobs run to completion with as little intervention as possible. Avoiding human 

interaction is critical to prevent errors in file input, compute-job staging, and post-processing. To 

address this problem, a software system has been developed to prepare each required simulation 

and submit each simulation to the compute-resource job manager. Basic graphics/visualizations 

and archiving are included to aid in QA/QC procedures and permanent storage of the computed 

results. 

This document describes the software production system developed for coupled 

ADCIRC/STWAVE simulations. The software manages the gathering and preparation of input 

files for the ADCIRC/STWAVE models, and writes the job control script that the user submits to 

the host machine job manager (e.g., LSF, PBS, LoadLeveler). The production management 

software is written in the scripting languages perl and bash. Technical details of the script 

procedures and operations are reported in a separate technical report. 

The primary requirements for the production system are independently verified and validated 

numerical models; tested and verified model grids; and runtime inputs to the models (forcing 

functions). Input files that provide initial conditions and forcing functions to the 

ADCIRC/STWAVE system are required to be pre-computed. This includes the storm realization 

(wind and pressure fields computed by the Ocean Weather PBL model) and the wave energy 

spectra for each storm (computed by the WAM model and interpolated onto the STWAVE 

model boundaries). 

Computer system 

The components of the production system are independent of the UNIX HPC computer 

system used, and can be set up on any system with accessible processing elements, large disk 

and storage capacity, and a job controller (Load Sharing Facility, Portable Batch System, 

LoadLeveler, etc). The scripting language perl and shell language bash must exist on the 

computer system (which is generally the case on Linux-based systems). Additional software 

packages (GNUplot, GMT) are used to facilitate post-simulation graphical analysis of the 

solutions. 

The production system has been developed on the ERDC Cray XT3 (Sapphire) and the 

University of Texas Dell cluster LoneStar. However, the system is easily portable to other highperformance computing facilities with minor changes to the scripts and job controller/manager 

configurations. This has been tested several times on the following system/job scheduler 

combinations: NAVO IBM P5 (Babbage)/LSF, LONI IBM P5 (Neptune)/LoadLeveler, Cray 

Research Cray XT3 (Salmon)/PBS. 

For the 2005, 2007, and 2010 (LACPR) simulations, the ERDC Cray XT3 Sapphire and 

University of Texas Dell cluster LoneStar were used for the computations. The only required 

changes to the scripts and setup are the specification of different job control parameters and a 

different home directory for the system. The basic computer system characteristics for Sapphire 

and LoneStar are shown in Table A. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-4

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table A: Basic configurations of the High-Performance Computer systems used for the 

LACPR storm-surge simulations. On 9 April, 2007, the ERDC Cray XT3 Sapphire 

returned to service after a major hardware upgrade and several related software 

changes. Its current configuration is indicated in parentheses in the table. 

System ERDC Cray XT3 Sapphire University of Texas Dell LoneStar

Operating System SUSE Linux & UNICOS/lc CentOS Linux 

Compute-Nodes AMD Opteron 2.6GHz single-core, singleprocessor (dual core, single-processor) 

Intel Xeon 5100 series 2.66GHz dual-core, 

dual-processor 

Number of 

Nodes/Processors 

4176/4176 (4096/8192) 1300/5200 

Parallel Filesystem (/work) Lustre Lustre 

Compute-Job Controller LSF (PBS) LSF 

Website www.erdc.hpc.mil www.tacc.utexas.edu 

Numerical Models in the Production System 

The “coupled” approach to the required storm-surge solutions uses the shallow-water finite 

element surge model ADCIRC (Luettich and Westerink, 2004) in conjunction with the steadystate spectral wave model STWAVE (Smith et al, 2001). Software versions are 46_48 and 

greater, and ep_110306 and greater, for ADCIRC and STWAVE respectively. Figure A shows 

the extents of the three models’ grids. The ADCIRC grids (boundary drawn in black) cover the 

western North Atlantic Ocean, with the only open boundary at 60 deg west. All ADICRC grids 

in this project use the same footprint and general coverage. Only the details in the Louisiana and 

surrounding regions differ. The Ocean Weather PBL model boundary is shown in red, covering 

the Gulf of Mexico and 0.05 degrees resolution. ADCIRC interpolates the PBL grid to the 

ADCIRC grid and pads the far-field extent with background values. The STWAVE grid 

boundaries (blue) are shown along the Louisiana, Alabama, and Mississippi coasts. The Lake 

Pontchartrain domain is barely visible at this scale. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-5

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure A. Footprints of the various model domains for the coupled ADCIRC/STWAVE system. ADCIRC, 

OWI/PBL, and STWAVE grid boundaries are shown in black, red, and blue respectively. 

Surge Model – ADCIRC 

The storm surge model ADCIRC is a state-of-the-art model that solves the generalized wavecontinuity equation on linear triangular elements. The depth-integrated (2D) implementation is 

used, where the water level and depth-averaged velocity are solved for at each triangle vertex 

(node), after complete specification of the initial conditions and time history of the boundary 

conditions (forcing). The finite element grids used in this project are the SL15 grid sequence for 

the 2005, 2007, and 2010 levee configurations. There are approximately 2.1 million horizontal 

nodes and 4.2 million elements in this sequence of grids. Ninety percent of the computational 

grid nodes are within the region shown in Figure B1. The computational domain is decomposed 

into 256 sub-domains, and each sub-domain problem is allocated to a separate computing 

processor. Inter-domain communication of boundary data is handled through MPI. Each solution 

is computed with a timestep of 1 second. A detail of the finite element grid is shown in 

Figure B2 for the New Orleans East area. 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure B1. ADCIRC bathymetry (in meters) for the SL15 grids along the northern Gulf of Mexico coast. In 

this image, the bathymetry is clipped at -100 meters for display purposes. Each of the SL15 

grids, for the 2005, 2007, and 2010 simulation sets has details that are not discernable at this 

scale shown. 

Figure B2. ADCIRC grid detail in the New Orleans East area. The solution is computed at the vertex of 

each triangle, and the thick black line is the model boundary. Weir boundary conditions are 

applied at the boundaries shown in this image. 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Wave model – STWAVE 

Five STWAVE grids are used in the current production system implementation covering the 

Louisiana, Mississippi and Alabama coastal region. (The “west” W grid is not used for the LA 

EAST simulations.) The boundaries of the domains are shown in Figure C. The spatial resolution 

of each STWAVE grid is 200 meters, and the temporal resolution is set to 30 minutes. The 

parallel STWAVE implementation is used in this system. In this version, each wave field 

snapshot is solved for on a separate processing element (cpu). The half-plane STWAVE is used 

for the South (S), Mississippi-Alabama (MS-AL), Southeast (SE) and West (W) grids. The fullplane STWAVE model is used for the Lake Pontchartrain domain. 

Figure C. Boundaries of the five STWAVE model domains and the PBL wind model northern extent (red) 

used in the current implementation of the production simulation system. The ADCIRC grid 

boundary is also shown (black) for one of the production grids. The Lake Pontchartrain (in 

green and not labeled) simulations use the full-plane STWAVE version; the other domains use 

the half-plane STWAVE. The “W” west grid is used in the LAWEST simulations, but not in the 

LAEAST simulations. The actual STWAVE grids to use for a specific simulation are a 

configurable part of the system. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Initial Condition to Storm Simulations (River Spinup) 

The initial condition for the storm-surge simulations is an ADCIRC solution that provides a 

background river discharge on the Mississippi and Atchafalaya Rivers. River inflow to the 

Mississippi River at Baton Rouge and to the Atchafalaya River at Simmesport is specified as a 

flux per unit width as defined by a wave radiation boundary. The radiation condition is based on 

the relationship between the normal flow and elevation at the boundary. The river condition is 

spun up for 2.0 simulation days, with forcing of normal flow specified at the head of the rivers. 

A two-day spin up period with a half-day hyperbolic ramping function is applied to the river 

boundary forcing prior to any additional model forcing. While the solution has not quite reached 

a steady state after this period, the subsequent wind-driven solutions continue this river spinup, 

and the river stages reach steady state over the next simulation day. The water elevations at the 

default sampling locations within the two rivers are shown in Figure D. This initial condition is a 

pre-computed requirement to each storm-driven simulation. 

Figure D. Water elevations along the Mississippi and Atchafalaya Rivers showing the spinup of the initial 

condition. This solution is driven by volume fluxes across the open boundary on the two rivers. 

The discharge is ramped up over the first 0.5 days. The maximum elevation reached at the top 

of the Mississippi River is 1.84 meters. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-9

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Wind Fields for Storms 

The primary inputs to the production system are the wind and pressure fields generated by 

application of the Ocean Weather Planetary Boundary Layer (PBL) model. The boundary of this 

domain is shown in Figure A1. Each storm/simulation must be represented by wind and pressure 

field files (WND/pre) in PBL format, ready to read by ADCIRC and as specified by the 

parameter NWS=+/-12,212 in the fort.15 file. For the LACPR storms, the WND/pre fields were 

pre-computed, external to the production system and assumed to exist as input to the system. An 

example of the WND/pre fields is shown in Figure E for LACPR storm number 042 (TPOC: Jay 

Ratcliff, USACE/MVN). 

Figure E. Wind (left) and pressure (right) fields for LACPR storm 042 at landfall. The storm track is 

shown with the dashed, red line. The fields are computed with the Ocean Weather Planetary 

Boundary Layer (PBL) model, and are used to drive the far-field wave model WAM and 

ADCIRC. The wind field information is passed to STWAVE through the ADCIRC global output 

wind file fort.74. 

The WND/pre files must be placed in the $PSHOME/winds directory to be detected by the 

production system. In the USACE/LACPR specification of the production system, the landfall 

date is YYYYMMDDHHMN, where the year is arbitrary and MMDDHHMN is 08010200 

(2AM on 01 August). This landfall date is conveyed to the production system scripts by 

specifying it in system control file (ProdSysDef.pm). It is also assumed that the storm center 

continues past landfall time for 24 hours. The pre-landfall time length is arbitrary. A graphic 

timeline is shown at the top of Figure F. 
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Figure F. Tracks used to define the 152 LACPR storms. Hurricane parameters vary along the tracks, and 

provide the input to the Ocean Weather PBL wind/pressure field model. This provides the 

storm forcing required for each production system simulation. 

There are 152 storms defined, each of which has a corresponding PBL WND/pre field. The 

storms are specified by variations of the hurricane parameters along the tracks shown in 

Figure F. The details of the storm parameter specification are described elsewhere in the 

technical documentation for this project. 

Wave Field Boundary Conditions for STWAVE 

All of the half-plane STWAVE grids (the full-plane STWAVE version is used for Lake 

Pontchartrain) require specification of the wave spectral characteristics on the open boundary. 

These boundary condition sets are pre-computed by running the WAM ocean wave model for 

each storm and extracting the wave energy spectra from the WAM solutions at the STWAVE 

open boundary node locations. The WAM-to-STWAVE procedure is applied for each of the 

storm wind/pressure fields, and the results are made available to the production system 

(WAM TPOC: Robert Jensen, USACE/ERDC-CHL). 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Solution Sampling Locations 

Each ADCIRC storm-surge simulation produces water level timeseries, both for the global 

model domain and for pre-defined station locations. In the current system implementation, the 

pre-defined sampling locations are shown in Figure G. There are three sets of points; for LACPR 

(blue), MSCIP (magenta), and IPET (red). The IPET locations correspond to reach locations as 

defined in the IPET risk model system definition files. The station output from each solution can 

be visualized using the post-processing code plotofort61.pl. This perl script generates a 

timeseries plot using GNUplot to be used for quick inspection of the fort.61 results. The 

timeseries plot for the ADCIRC3 (step 4) part of the process is shown in Figure G for storm 042. 

Figure G. Left) ADCIRC solution sampling points that define the station output in the fort.61 file. The 

global model solution is interpolated to these locations and output for visualization and 

analysis. Right) Hydrograph plot generated by plotfort61.pl, for the ADCIRC3 step of the 

production system, for LACPR storm 042. 

Solution Procedure (Process Management) 

Each simulation is performed in four computational parts, as shown in the “Production 

Steps” part of Figure H. The initial conditions (River Spinup), PBL wind and pressure fields, and 

the wave energy spectra boundary conditions (as computed by WAM) are all required inputs to 

the system. 

1. ADCIRC1: Each simulation is started from the River Spinup (ADCIRC0). ADCIRC is 

run from the start of the PBL wind field to 24 hours prior to landfall of the storm 

(ADCIRC1, River+Winds). The model state is output to disk to provide initial conditions 

for continuation of the simulation (step 2), and for the subsequent rerun of step 2 that 

includes wave radiation stresses from STWAVE (step 3). 

2. ADCIRC2: The ADCIRC1 solution (River+Winds) is then continued to the termination 

of the PBL wind/pressure fields. The ADCIRC2 global water level (fort.63) and wind 

field (fort.74) are output for QA/QC as well as to provide to STWAVE as input. 

3. STWAVE: The ADCIRC2 solution in step 2 is interpolated to the STWAVE domains, 

and each STWAVE domain is executed to generate wave radiation stress gradients for 

input back to ADCIRC in step 4. 
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4. ADCIRC3: ADCIRC2 is re-run over the same time period as in step 2, but including the 

wave radiation stress gradient computed by STWAVE and interpolated onto the 

ADCIRC grid. This is the River+Winds+RadStress solution and is also referred to as the 

ADCIRC+STWAVE step. 

Figure H. Timeline for Production System simulations. The timeline for the wind forcing (PBL) is shown at 

the top. PBL winds and pressures are used to drive both the WAM wave model and the 

ADCIRC/STWAVE simulations. Specification of the PBL and WAM fields is a required input to 

the Production System. 

These four steps are coordinated by a perl script (prep_prod_sim.pl) that organizes the 

directory structure for each run, gathers the needed input files (grids, initial conditions), edits the 

files for the specifics of the current simulation, and writes a job control file for the highperformance computer system job controller (PBS on the ERDC Cray Sapphire and LSF on the 

Uni. of Texas Lonestar Dell cluster). The data file flow controlled by this script is shown in 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-13

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure I. To stage a specific simulation, this script is executed with the storm number specified 

on the command line. 

Figure I. Data flow controlled by the production system script prep_prod_sim.pl. Pre-computed inputs 

are shown at the top with red arrows. The middle row of boxes corresponds to the four steps, 

where the STWAVE step includes interpolation from ADCIRC to the STWAVE grids, the 

STWAVE simulations, and the interpolation back to ADCIRC. ADCIRC2+STWAVE is the 

“ADCIRC3” step. Blue lines indicate the ADCIRC hotstart file process. Parameters that are the 

same for all simulations are defined in the ProdSysDef.pm file and read by prep_prod_sim.pl at 

runtime. 

Katrina Simulation with PBL Wind Field 

A comparison of the observed high-water marks (HWMs) to ADCIRC+STWAVE solutions 

is shown in Figure J, for the surge that accompanied hurricane Katrina. Two ADCIRC solutions 

are shown. 

The verification solution, forced by the Ocean Weather 95% best-estimate winds, is 

compared to the observed HWMs (in blue). The overall agreement is very good, with the bestestimate winds solution explaining 94% of the observed HWM variance (r2

=.94, Se=1.5). The 

ADCIRC3 solution forced by the Ocean Weather PBL wind and pressure field is shown in red. 

The agreement is also very good (r2

=.88, Se=2.1). 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-14

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

For each solution, the best-fit line is shown, colored the same as the data points, as well as 

the 95% confidence region about the data. 

Figure J. Comparison of observed High Water Marks (HWM) for hurricane Katrina, and the Production 

System Katrina simulation using the PBL wind and pressure fields (red). Also shown is the 

comparison between the observed HWMs and the verification solution forced by the Ocean 

Weather 95% wind and pressure field (blue). Observations are on the abscissa, and the 

ADCIRC3 values are on the ordinate. 

References 

Luettich, R., and J. Westerink. Formulation and Numerical Implementation of the 2D/3D ADCIRC, Finite 

Element Model Version 44.XX, 2004. http://adcirc.org/adcirc_theory_2004_12_08.pdf
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-15

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Smith, J., A. Sherlock, and D. Resio. STWAVE: Steady-State Spectral Wave Model User's Manual for 

STWAVE Version 3.0, Tech Report ERDC/CHL SR-01-1, 2001. 

http://chl.erdc.usace.army.mil/Media/2/4/4/erdc-chl-sr-01-11.pdf
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Appendix 8-1 

Acronyms 

ADCIRC ADvanced CIRCulation model for shallow seas 

CHL Coastal Hydraulics Laboratory 

ERDC Engineer Research and Development Center 

GMT Generic Mapping Tools; open-source plotting package used for post-processing 

graphics generation 

HWM High-water mark 

HPC High-performance computing 

I/O Computer input and output 

IPET Interagency Performance Evaluation Task force 

LACPR Louisiana Coastal Protection and Restoration project 

LONI Louisiana Optical Network Initiative 

LSF Platform Computing’s Load Sharing Facility (LSF) job scheduler 

MPI Message-passing interface; parallel programming model used by ADCIRC and 

STWAVE 

NAVO NAVal Oceanographic computing facility 

PBL Planetary boundary layer model of Ocean Weather Inc. 

PBS Portable Batch System job scheduler 

QA/QC Quality assurance/quality control 

STWAVE STeady-state spectral WAVE model 

WAM NOAA large-scale wave model 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-17

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Appendix 8-2 (R2007) 

(May 27, 2007) 

White Paper on Estimating Hurricane 

Inundation Probabilities 

Written by: Donald T. Resio, Senior Scientist, U.S. Army Corps of Engineers, ERDC-CHL 

Incorporating contributions, discussions, data, and comments by: (in alphabetic order) 

Stanley J. Boc (ERDC-CHL), Leon Borgman (private consultant), Vincent J. Cardone, 

(Oceanweather, Inc), Andrew Cox (Oceanweather, Inc.), William R. Dally (Surfbreak 

Engineering), Robert G. Dean (U. of Florida), David Divoky (Watershed Concepts), Emily 

Hirsh (FEMA), Jennifer L. Irish (Texas A&M University), David Levinson (NOAA’s National 

Climatic Data Center), Alan Niedoroda (URS Corporation), Mark D. Powell (NOAA’s 

Hurricane Research Division, AOML), Jay J. Ratcliff (USACE-MVN), Vann Stutts (USACEMVN), Joseph Suhada (URS Corporation), Gabriel R. Toro (Risk Engineering), Peter J. 

Vickery (Applied Research Associates), and Joannes Westerink (U. of Notre Dame) 

Introduction 

Over the last several months, a team of Corps of Engineers, FEMA, NOAA, private sector 

and academic researchers have been working toward the definition of a new system for 

estimating hurricane inundation probabilities. This White Paper is an attempt to capture the 

findings and recommendations of this group into a single document. 

At least five methods have been applied in past studies of environmental extremes due to 

hurricanes in the United States: 

1. Formulation of design storm events 

2. Estimates based only on historical storms 

3. The empirical simulation technique (EST) 

4. The joint probability method (JPM) and 

5. The Empirical Track Model 

In each method, it is important to understand that two different statistical measures are 

required to characterize the expected extremes over an interval of time. The first of these is the 

measure of the expected values of the distribution and the second is a measure of expected 

dispersion around these central values. 

All of the methods referenced above have different strengths and weaknesses for various 

applications. To help understand these, a brief discussion of conventional applications of these 

methods in the past will be given before moving on to a discussion of the approach 

recommended for future hurricane surge applications. This will be followed by sections that treat 

the following: 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-18

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

1. Potential extensions of probabilistic methods to future hurricane surges, 

2. The JPM with Optimal Sampling (JPM-OS); and 

3. A computational methodology for effective simulation of storm surges for hurricane 

inundation studies. 

Formulation of Design Storm Events 

An example of this approach applied to coastal inundation is the Standard Project Hurricane. 

This approach was adopted by the Corp of Engineers in the 1960’s to estimate potential surge 

hazards along many U.S. coastlines. Due to the paucity of data, it would have been very difficult, 

if indeed possible, to investigate detailed characteristics of landfalling hurricanes prior to 1960; 

consequently, the Corps requested that NOAA prepare an estimate of a storm with characteristics 

that were expected relatively infrequently within some stretch of coastline. Unfortunately, the 

period prior to 1960 (the input to the statistical analyses performed by NOAA) was a period of 

relatively low hurricane activity in the Gulf of Mexico; consequently, the SPH, as specified in 

those earlier studies, is not representative of the characteristics of extreme storms that have 

occurred in the Gulf since 1960. 

In the past, the design storm approach has typically utilized a single storm to characterize 

environmental factors for design at a given location. This effectively reduces the number of 

degrees of freedom in storm behavior to one parameter, typically the intensity of the storm. All 

other storm parameters (for example: storm size, forward storm speed, and track location) are 

deterministically related to storm intensity. The major problem with this is that, if a second 

factor (such as angle of storm approach to the coast, storm duration or river stage) significantly 

affects design conditions and/or considerations, the design storm approach cannot accurately 

capture all aspects of the storm that affect the design. An extrapolation of the single design storm 

concept is to define a small set of storms with some range of additional parameters considered. 

This defines a set of storms that can be used to examine various design alternatives in an 

efficient manner, while retaining some additional degrees of freedom within the analysis. 

Because of their use in this context, these storms are often termed “screening storms” rather than 

design storms. 

Estimates Based only on Historical Storms 

If at least one sample from a population of interest occurs within each year, it might be 

possible to apply statistical methods that utilize annual maxima to formulate the stage-frequency 

curves at a site. However, it is clear that such a situation cannot be well characterized by 

conventional asymptotic methods which assume many samples in a population occur annually. 

Since hurricanes are both relatively infrequent and relatively small in terms of the amount of 

coastline affected by these storms each year, the frequency of storms affecting a site is typically 

significantly less than 1 storm per year. As an example of this, Figure 1 provides an estimate for 

the frequency of severe hurricanes (greater than Category 2 intensity) per 1-degree by 1-degree 

(area) per year for the Gulf of Mexico. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-19

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Another sampling problem related to the use of historical storms for specifying extremes is 

the tendency for intense storms to behave differently than weaker storms. To compensate for this 

sampling inhomogeneity, many oil-industry groups have adopted the “Peaks Over Threshold,” or 

POT method for estimating extremes. In this approach, only storms above some threshold value 

are considered within a statistical analysis of extremes. By screening small storms from the 

analyzed sample, the effect of small storms on the parameters of fitted distributions (e.g. the 

parameters of the Gumbel, Frechet, Weibull, Lognormal, Log Pearson or other distributions of 

choice) is minimized. This approach is inherently parametric due to the need to assume/specify a 

distribution (or class of distributions such as the Generalized Extreme Value method). 

Another potentially more serious problem with the reliance only on historical storms for 

estimating coastal inundation is related to the small sample of storms at any site. As can be seen 

from Figure 1, even in the vicinity of New Orleans, the frequency of major storms passing within 

a given 60 nautical mile region (1-degree) is only about once every 20-25 years. Since we have 

limited records beyond the middle of the 20th century and since the frequency of storms in the 

latter half of the century may be markedly different than the frequency of storms in the early part 

of the century, the historical record for a direct hit (taken here for simplicity as ±30nm of a site) 

would include on the average only about 2-3 storms. Given this small potential sample size, 

sampling variability can lead to very unrealistic variations in storm frequency and intensities 

along the coast. For this reason, methods for estimating coastal inundation based solely on 

historical analyses should not be used for coastal hazard assessment. This point is also pertinent 

to arguments that the use of historical records for extremal estimation is also difficult to justify in 

many other, non-coastal applications. 

The Empirical Simulation Technique (EST) 

Statistics of Expected Extremes 

Conventionally, the EST was used as an approach for estimating the expected extremal 

distribution for surges based on a variation of the “historical record” approach (Borgman et al., 

1992; Scheffner et al. 1993). In this approach a set of storms above some threshold affecting a 

particular area are hindcast, similar to the approach used within the POT method. The primary 

difference is that typically the largest 1 or 2 storms is re-run over a number of track variations in 

order to distribute the effects of the storm over a wider area. Unfortunately, the rules for this redistribution of tracks were developed only after some preliminary applications of this approach 

and tend to be somewhat arbitrary. Results from computer simulations at each point of interest 

are then ranked and assigned a cumulative distribution value via a formula that links the rank to 

the cumulative distribution function, F(x), which is be abbreviated as CDF is subsequent 

discussions, 

( ) 1

m F x

N = +

 (1) 

where m is the rank of the storm (with m=1 as the smallest) and N is the total sample number. 

This is converted to a measure of recurrence via the use of a Poisson frequency parameter, λ
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1 ( ) [1 ( )]

T x

λ F x = − (2) 

where T(x) is the expected return period for x. For example, if we hindcast 50 storms from an 

interval of 100 years, the value of λ would be 0.5; and the estimates of T(x) would span the 

range from slightly over 100 years to slightly over 2 years. 

In the interior of the ranked points, the EST assumes that the best estimate of the expected 

distribution is the sample itself and does not fit any parametric distribution to the central 

distribution in order to obtain a “smoothed” distribution. Thus, within this interior range, the 

distribution is nonparametric. If one is interested in return intervals outside of the range covered 

by the ranked sample points, it is necessary to invoke some sort of empirical (parametric) 

function for extrapolation. This is a severe limitation of the conventionally applied EST, since it 

restricts the non-parametric estimates to approximately the number of years covered by the 

hindcasts, just as in any other method that relies only on historical storms. In fact, the 

conventional EST, as used in most past studies, is very similar to the so-called peaks over 

threshold (POT) method based on historical hindcasts, with three notable exceptions discussed 

below. 

First, the POT method typically uses a parametric fit in the interior region of the distribution, 

rather than just in the region beyond the largest storms. It can be shown via Monte Carlo 

simulations that, if the entire sample is drawn from a single parent distribution, the POT method 

would provide a more stable estimate of the actual distribution in the interior than would the 

EST. However, in many situations in nature, the interior distribution can consist of samples from 

several different parent populations. An example of multiple populations would be the case of 

storm surges due to “direct hits” of hurricanes vs. storm surges due to bypassing storms. Surges 

from these two sources could have relatively different probability characteristics at a fixed 

coastal location, as shown in Figure 2. Consequently, the use of nonparametric estimates within 

the central portion of the distribution seems more appropriate for most coastal flooding 

applications. 

Second, the POT method uses a “fit” to the entire sample to extrapolate into longer return 

periods. In this case, the problem of mixed populations could potentially introduce serious 

problems into extrapolations, if the more that one population persists into the region above the 

“cut-off” threshold. However, many POT methods carefully choose the threshold value to try to 

restrict the sample to a single population. In some situations this is possible and in others this 

may not be possible due to lack of sample size (i.e. there may only be 1 or two direct “hits” by 

hurricanes). On the other hand, the EST uses an empirical “spline” fit to some small number of 

points near the high end of the distribution combined with a set of secondary restrictions that 

tend to limit excessive curvature in the extremes. Extrapolations in such situations tends to be 

subject to considerable judgment, since there is no underlying distribution used as the basis for 

estimating values larger than those included within the sample. 

Third, in hurricane surge studies with the EST, it was observed that a single intense storm 

could introduce an anomalously large value into a small spatial region, while other areas had 

much lower maximum surge values. Since there is no reason to believe that all future hurricanes 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

would strike only points included in the historical data set, some applications of the EST 

introduced a set of hypothetical storms intended to distribute the effects of large storms 

throughout the region being studied. In this approach, the largest storm was usually assigned a 

number of offset tracks and the probability of that storm was distributed over that set of tracks. 

However, each of these “cloned” storms had exactly the same characteristics as the original 

storm. In this approach, the extent of the offsets was rather arbitrary, in spite of the fact that the 

degree of smoothing (distance over which the storm track is replicated) can significantly affect 

estimated local extremes. Also, only characteristics of storms that fell within the area being 

analyzed were included in an analysis. Thus, a storm such as Katrina (a large, very intense 

storm) is difficult to reliably predict from a local sample that does not contain this type of storm 

within its sample. 

Both the POT method and the EST use various statistical methods to incorporate tidal effects 

into their estimates of water levels at the coast. In a sense, these could be considered as 

hypothetical storms, since they represent storms that struck the coast at a different tidal phase. 

From this discussion, it can be seen that a major problem with the conventional EST is the 

same as encountered in any approach to hurricane flooding based on local data only: lack of 

sample. The relatively small spatial extent of hurricanes combined with the small number of 

storms affecting a given section of coast during the period of reliable records and the apparent 

existence of long term cycles and trends makes the sampling variability in historical records very 

large along U.S. coastlines. This same variability affects all estimates of coastal flooding since 

they all rely on historical data; however, this variability can be greatly exacerbated when only 

local data is considered. It is perhaps fair to say that the application of the EST requires more 

experience and judgment than an application of the JPM. 

Measures of Variability in Expected Extremes 

Since the EST does not contain an underlying theoretical distribution, no theoretical 

estimates of variability (confidence limits) are possible. Instead, the EST uses a “re-sampling” 

method to obtain estimates of expected variations in extremes over an interval of time. This 

method, rather than its methodology for estimating expected extremes, is the strength of the EST. 

The combination of the re-sampling with the nonparametric distribution provides a sound basis 

for estimating the variability of the estimated extremes. 

The Joint Probability Method (JPM) 

Statistics of Expected Extremes 

The JPM was developed in the 1970’s (Myers, 1975; Ho and Meyers, 1975) and 

subsequently extended by a number of investigators (Schwerdt et al., 1979; Ho et al., 1987) in an 

attempt to circumvent problems related to limited historical records. In this approach, 

information characterizing a small set of storm parameters was analyzed from a relatively broad 

geographic area. In applications of this method in the 1970’s and 1980’s, the JPM assumed that 

storm characteristics were constant along the entire section of coast from which the sample was 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

drawn. Recent analyses suggest that this assumption is inconsistent with the actual distribution 

of hurricanes within the Gulf of Mexico. 

The JPM used a set of parameters, including 1) central pressure, 2) radius of maximum wind 

speed, 3) storm forward speed, 4) storm landfall location, and 5) the angle of the storm track 

relative to the coast, to generate parametric wind fields. Furthermore, initial applications of the 

JPM assumed that the values of these five parameters varied only slowly in storms approaching 

the coast; therefore, the values of these parameters at landfall could be used to estimate the surge 

at the coast. Recent data shows that this is not a good assumption (Figure 3). Kimball (2006) has 

shown that such decay is consistent with the intrusion of dry air into a hurricane during its 

approach to land. Other mechanisms for decay might include lack of energy production from 

parts of the hurricane already over land and increased drag in these areas. In any event, the 

evidence appears rather convincing that major hurricanes begin to decay before they make 

landfall, rather than only after landfall as previously assumed. 

The conventional JPM used computer simulations of straight-line tracks with constant 

parametric wind fields to define the maximum surge value for selected combinations of the basic 

five storm parameters. Each of these maximum values was associated with a probability 

max

max

( , , , ,)

where

 is the central pressure,

 is the radius of maximum wind speed

 is the forward velocity of the storm

 is the angle of the track relative to the coast at landfall

 is the 

p fl

p

f

l

pc R v x

c

R

v

x

θ

θ

distance between the point of interest and the landfall location 

These probabilities were treated as discrete increments and the CDF was defined as 

() | F ijklm ijklm x px x = < ∑ (3) 

where the subscripts denote the indices of the 5 parameters used to characterize the hurricanes. 

Similar to the EST, this method is nonparametric; however, the conventional JPM included a 

range of parameter combinations that typically made extrapolation beyond the range of 

simulations unnecessary. This is an advantage over the conventional EST, since it removes the 

need to assume a particular parametric form for the CDF in critical ranges of values. 

Another potential advantage of the JPM over methods which depend heavily on historical 

storms was that the JPM considered storms that might happen; whereas, the EST considered only 

storms that did happen. Assuming that, for the purpose of surge generation, storm characteristics 

can be represented adequately by the set of parameters used, it is possible to construct a Katrinalike storm (high intensity combined with large size) even if one has not happened previously. 
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Likewise, it is possible to interpolate between re-curved storms such as Opal and Wilma to 

understand possible hurricane impacts in the Tampa area, even though neither of these storms 

produced significant surges in the Tampa area. 

Perhaps the biggest controversy in JPM applications during the 1970’s and 1980’s centered 

on the definition of this 5-dimension joint-probability function. The lack of data on historical 

storms prior to 1950 made it very difficult to derive representative distributions, even for 

extended sections of coast. For example, information on storm size (radius of maximum wind 

speed was lacking for most historical storms; consequently, a statistical estimate of max r (as a 

function of latitude and central pressure) was frequently substituted for actual values in the 

probability distribution. One wind field factor not considered in early JPM applications was the 

variable peakedness of hurricane wind fields. This term is represented in terms of the Holland B 

parameter in recent hurricane wind models and will be discussed in a subsequent section of this 

white paper. 

One point of interest that should not be lost here is the importance of capturing the mean 

statistical behavior of any time-varying properties used in JPM applications. For example, surges 

derived from previous JPM applications, under the assumption of that storm characteristics near 

the coast were constant, may have been biased low, since they were based on statistics at 

landfall. Since storms are consistently more intense off the coast (as shown in Figure 3), the 

modeled offshore storms are less intense than the actual offshore storms, under this assumption. 

Of course, some calibration was performed in these studies, so this might have been somewhat 

accounted for via calibration procedures; however, calibration tends to somewhat storm specific, 

so such calibration could still leave considerable residual bias in the final results. 

JPM Measures of Variability in Expected Extremes 

Most applications of the JPM only considered the definition of the mean CDF from the 

simulations, and little attention was paid to quantifying the dispersion (uncertainty) of what 

could happen within a particular time sample. This is potentially a major shortfall in the JPM as 

it was originally applied. 

The Empirical Track Model 

Vickery et al. (2000) presents a method for modeling hurricane risk in the United States. 

This method has been adopted for the development of design wind speed maps within the U.S. 

(American National Standards Institute (ANSI), ASCE 1990, 1996). This method uses a Monte 

Carlo approach to sample from empirically derived probability and joint probability 

distributions. The central pressure is modeled stochastically as a function of sea surface 

temperature along with storm heading, storm size, storm speed, and the Holland B parameter. 

This method has been validated for several regions along U.S. coastlines and provides a rational 

means for examining hurricane risks associated with geographically distributed systems such as 

transmission lines and insurance portfolios. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-24

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A key requirement for the application of the Empirical Track Model within its Monte Carlo 

framework is the ability to efficiently execute storms over many, many years (20,000 years in the 

Vickery et al. (2000) application). Whereas this is not too demanding for an efficiently written 

PBL wind model, it is well beyond the range of possibility in large, high-resolution ocean and 

coastal response models (wave models and surge models). For this reason, the Empirical Track 

Model was not considered for application to coastal inundation; however, it provides an 

excellent source for validating the statistical characteristics of the winds used for inundation 

modeling. 

Potential Extensions of Probabilistic Methods to Future 

Hurricane Surges 

As noted in the earlier section on the design storm approach, the design storm approach 

suppresses much of the real variability in storm characteristics and should be used only for 

screening alternatives and not for final design of critical coastal structures. However, it is 

sometimes informative to examine the relative return period of specific historical storms. 

Appendix A provides an estimation method along these lines which estimates the return periods 

for both Hurricanes Katrina and Rita (found to be 397 and 89 years, respectively). 

It should be clear from the previous section that both the EST and the JPM in older 

applications suffered from a paucity of historical data. Also, older applications did not include 

many of the modeling advances that are now regarded as necessary for accurate simulations, for 

example: inclusion of numerically simulated wave set-up, use of detailed grids to capture highresolution bathymetric effects, and the application of improved near-coast meteorological models 

for hurricane evolution and wind-field behavior. Many of these effects were recognized and 

improved methods for treating them were developed during the forensic study of Hurricane 

Katrina by the Interagency Performance Evaluation Task Force (IPET). 

Recognition that waves can play a substantial role in coastal surge levels introduces a key 

difference between older JPM applications and the applications being considered today. In the 

older methods, the assumption of constant parameters was intended to span an interval of about 

12 – 24 hours, essentially the primary period for direct wind forcing of coastal surges in 

historical hurricanes. Waves contribute to coastal surges primarily in nearshore areas; however 

they are generated over a span of days during the approach of the storm to land. Thus, 

consideration of wave contributions to surges requires some knowledge of the storm behavior 1-

4 days prior to landfall. Treatment of hurricanes in terms of straight tracks with constant size and 

intensity over such a period is a bad assumption, since such tracks would not retain the wavegeneration characteristics produced by the curved tracks within the historical record. 

This brings us to a crucial point in considering what statistical approach should be used in 

future applications, i.e. whether to use an approach based strictly on re-sampling historical 

storms (essentially a modification/extension of the EST) or an approach based on parameterized 

wind fields over longer tracks (essentially a modification/extension of the JPM). On one hand, it 

is clear that each historical storm has many, many factors which vary throughout its history 

which influence hurricane wind fields. For example, eye-wall replacement cycles, interactions 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-25

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with large-scale wind systems, asymmetries within the eye-wall, and complicated track 

curvatures can all create significant perturbations within the structure of hurricane wind fields. It 

is not clear however that such modifications to the winds affect the total wave generation process 

so much, since the nature of this process integrates the wind input over several 10’s of hours. In 

fact, numerical experimentation with hurricane Katrina showed that all of the versions of the 

wind field (PBL alone, HWIND, and the most recent Oceanweather version) created minimal 

variations in the wave field (in the range of 5-7% differences in wave height) in the area off of 

Mississippi. 

In a sense, the critical factor that must be considered here is the number of primary 

dimensions required for representing wind fields with sufficient accuracy that they provide 

reasonable, relatively unbiased skill when used to drive coastal wave and surge models. For the 

case of extratropical storms, there is no known simple set of parameters that meets this criterion 

and some extension of the EST or POT method may be the suitable choice for such applications. 

For the case of hurricanes, dynamic models of hurricane wind fields (Thompson and Cardone, 

1996; Vickery et al., 2000) can be shown to capture a substantial portion of the wind field 

structure, when driven with the parameters listed above plus the so-called “Holland B” (Holland, 

1980) parameter. 

Each hurricane will tend to exhibit some degree of deviation from the theoretical PBL-model 

estimates. At any fixed time, such deviations could be produced by strong storm asymmetries, 

variations in Rmax around the storm, enhanced spiral bands, etc. Hence, a “best-estimate” wind 

field crafted by experts to assimilate all the observations in a given hurricane will typically 

represent the details of that particular storm much more faithfully than possible via a 

parameterized theoretical model. Such wind fields today are produced primarily by Mark Powell 

and others at the Hurricane Research Division of the National Hurricane Center of NOAA or by 

analysts at Oceanweather, Inc. These wind fields are absolutely essential for advancing our 

understanding of hurricane winds relative to wave and surge forcing in offshore and coastal 

areas. 

It is obvious that “best-estimate” wind fields contain an extremely large number of degrees 

of freedom in their formulation. Given the relatively small number of historical hurricanes, it is 

unlikely that we can understand/quantify the probabilistic nature of all the interrelated detailed 

factors creating these deviations. If these details were absolutely critical to coastal wave and 

surge estimation, we would be able to represent a past hurricane very accurately but would know 

little about the probability of future hurricanes unless we retained the same number of the 

degrees of freedom, including their expected variability in estimates of future storm surge and 

wave estimates. To demonstrate this point, in the definition of surge probabilities via numerical 

models, we are considering a relationship of the form 

12 12 12 ( ) ... ( , ,..., ) [ ( , ,..., ) ] ... p η δη = Ψ− ∫ ∫ p x x x x x x dx dx dx n nn (4) 

where max η (= for each individual storm at a fixed spatial location) η is the storm surge level, 

δ[.] is the Dirac delta function and 1 2 ( , ,..., ) n Ψ x x x is a numerical model or system of models 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-26

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that operate on the set of parameters ( 1 2 , ,..., ) n x x x to provide an estimate of the surge level at a 

fix location. This can be directly integrated to yield the CDF for surge levels 

12 12 12 ( ) ... ( , ,..., ) [ ( , ,..., )] ... F p x x x H x x x dx dx dx η η = −Ψ ∫ ∫ n nn (5) 

where H[.] is the Heaviside function. If we retained a sufficient number of degrees of freedom 

to resolve the wind fields exactly, if our numerical codes were also “exact,” and if our 

specification of the joint probability function 1 2 ( , ,..., ) p n xx x were known exactly, we could treat 

this equation as an exact integral for the CDF, with no uncertainty in its expected value. The 

sampling variability could then be estimated by re-sampling methods along the lines of the EST. 

The CDF integral (equation 5) shows that the number of dimensions required for an exact 

representation of the surge CDF must equal the number of degrees of freedom contained within 

the system (for practical purposes determined by the number of degrees of freedom contained in 

the wind fields). Since we recognize that all wind fields, wave models, and surge models remain 

inexact and that our estimates of joint probabilities are greatly hampered by lack of sample size, 

it is clear that the actual representation of this integral should be written as 

12 12 12 ( ) ... ( , ,..., , ) [ ( , ,..., ) ] ... F p x x x H x x x dx dx dx d η n nn = −Ψ + εη ε ε ∫ ∫ (6) 

where ε is an “error” term due to wind field deficiencies, model deficiencies, unresolved scales, 

etc. In this form, we see that there is a trade-off between modeling accuracy and the magnitude 

of the error term, ε . There is also a similar trade-off between errors/uncertainties in the 

probability estimates and the overall accuracy in estimates of the surge CDF. These errors will 

increase substantially if we attempt to split a small sample (for example the historical hurricane 

record in the Gulf of Mexico) into information for too many dimensions. Following this 

reasoning, it seems advisable to limit the number of parameters considered in the JPM 

probability integral and to include an approximation for all of the neglected terms within the 

error term, ε . As noted previously, PBL models provide a relatively accurate representation of 

the broad-scale structure within hurricanes. Furthermore, wind fields from PBL models have a 

very long history of providing accurate wave estimates in Gulf of Mexico hurricanes (Cardone et 

al., 1976). Consequently, the logical choice appears to be to limit the number of dimensions in 

the JPM integral to the number of parameters contained within such PBL models 

12 12 ( ) ... ( , , , , , ) ( ) [ ( , ,..., ) ] ... F p c R v x B p H x x x dx dx dx d η = −Ψ + p pfl n n θ εη ε ε ∫ ∫ (7) 

where the error term has been separated from the rest of the probability distribution. In this form, 

the “error” term allows us to include additional effects on water levels, such as tides (albeit in an 

uncoupled, linear fashion). Also, in this equation, we have replace Rmax with Rp , since the latter 

term is used in the PBL model selected for application here (see Appendix B) for details. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-27

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During the last several months considerable effort has gone into re-analyzing hurricane 

characteristics and hurricane wind fields. One of the significant findings of this effort is that the 

Holland B parameter in mature storms within the Gulf of Mexico tended to fall within the range 

of 0.9 – 1.6. Furthermore, numerical sensitivity tests of both wind fields and coastal surges 

suggest that the adoption of a constant value of 1.27 for storms centered more than 90 nm from 

the coast provided a reasonable first approximation to both the wind fields and the surges. Thus, 

if we add the effects of B-variations into the “error” term, we can reduce the CDF equation to 

12 12 ( ) ... ( , , , , ) ( | ) [ ( , ,..., ) ] ... F p c R v x p H x x x dx dx dx d η = −Ψ + p pfl n n θ εη η ε ε ∫ ∫ (8) 

In this form, ε is considered to include, at a minimum, the following terms: 

1. tides, 

2. random variations in B, 

3. track variations not captured in storm set, 

4. model errors (including errors in bathymetry, errors in model physics, etc.), and 

5. errors in wind fields due to neglect of variations not included in the PBL winds. 

It is evident that we can only approximate the overall distribution of ε from ancillary 

information on errors in comparisons to High Water Marks and comparisons of results from runs 

with the “best-estimate” wind fields and PBL wind fields. Tides can be factored into this analysis 

assuming linear superposition, with some degree of error introduced. Based on the best available 

approximations to all of these terms, assuming that all the “error” contributions are independent, 

and a loose application of the Central Limit Theorem, we will assume that the “error” term can 

be represented as a Gaussian distribution with a mean of zero (assuming that the model suite is 

calibrated to this condition) and a standard deviation equal to some percentage of the modeled 

surge. 

If we were to try to extend the EST to include shifting actual historical storms from one 

landfall site to another, we would implicitly be holding all the “details” constant by assuming 

exactly the same storm will occur in the future at different locations. This constrains all future 

hurricanes to have the same detailed characteristics as the single historical storm upon which it is 

based. It seems more logical to approach this in two stages as described above, an initial stage to 

capture the broad-scale wind characteristics and a second stage to understand/quantify the impact 

of deviations around these broad-scale winds on coastal surges. In this way, the probability 

analysis can be kept within a sufficiently small number of dimensions to allow reasonable 

approximation from historical records. 

In our own work investigating modifying the EST, we began by making the same assumption 

as the older JPM application, i.e. that we could take some section of coast and treat it at though it 

were homogeneous with respect to expected storm parameters. With this assumption, we felt that 

it would be appropriate to move a storm some distance along the coast without affecting its 

characteristics. Unfortunately, when we tested this approach, we saw that historical tracks did 

not translate well spatially. We found that all of the large storms affecting U.S. Gulf of Mexico 

coastlines entered the Gulf through either the gap between Cuba and the Yucatan Peninsula or 

through the southern Florida to Cuba area. A simple geometric displacement of a particular 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-28

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storm track made that track intersect with land in areas that are not suitable for their origin. For 

example, such a displacement to allow a storm such as Hurricane Opal to strike the western 

portion of Louisiana would intersect with land at a point slightly south of the border between the 

United States and Mexico. It is essentially impossible for such a track to generate a storm of 

Opal’s strength. Another example would be the simple translation of Katrina’s track to a point 

farther east. It is highly unlikely that such a track would be able to support a storm of Katrina’s 

size and intensity. Thus, any concept that involves a simple geometric translation of a historical 

storm track to a position very far from where it actually occurred was found to be very naïve, at 

least in a meteorological/climatological sense. Still another example of problems associated with 

the use of a detailed track would be the simplistic shifting of a track with a loop in it, such as 

Hurricane Elena along the panhandle of Florida. The implication of using this exact track and 

associated wind field is that all future storms of this type will exhibit exactly this same loop and 

associated intensity, size, asymmetry, and other detailed characteristics during its approach to 

land. The probability of this actually occurring is very near zero. 

An additional problem in our initial attempt to modify the EST for storm surge application 

relates to the definition of storm probability. We found that, due to the geographic constraints of 

entry points for intense storms into the Gulf, it was a very bad assumption to treat any extended 

section of coast as though it were a homogeneous area in terms of expected hurricane 

characteristics. Instead, several independent analyses have shown that the statistical properties of 

hurricanes vary continuously and substantially throughout the Gulf of Mexico. This means that it 

is not advisable to shift a storm track from one section of coast to another and treat it as an 

equivalent sample to a storm that actually occurred in that section of coast. A related problem in 

attempting to modify the EST for hurricane surge applications is assigning the probability of 

exceedance of some characteristic in a particular storm. For example, most critical reviews 

conducted recently have advocated limiting the period of record to be used in climatological 

estimates of hurricane characteristics to a period from no earlier than the mid-1940’s to the 

present. Thus, if we adopt the conventional plotting position, Katrina will represent somewhere 

in the neighborhood of a 60 year event. We can probably increase this a little by spreading 

Katrina over some section of coast; but as noted above, this is not a simple exercise. The degree 

of spreading, the reasonableness of the tracks, and the spatial variation of statistical properties all 

make this a procedure that should not be trivialized. 

The final point worth noting is that prior to 2005, there was no Katrina in the historical 

record. Straightforward estimates of surges using the EST method prior to Katrina suggest that 

such a surge only occurs once every 1200 years, or so. Most analyses suggest that this is a 

substantial overestimation of the return period for such a surge. The problem here is that, prior to 

Katrina, no historical storm combined both intensity and size in a fashion equivalent to Katrina. 

Thus, Katrina was outside the sample range of the EST. Although this particular deficiency in 

the sample has been remedied for at least some of the Gulf coast (since Katrina has now occurred 

there), one must ask what other gaps in the sample of historical storms might exist. Given this 

point along with all the other discussion points in this section, it was decided that it would be 

more straightforward to modify the JPM for hurricane surge applications than to modify the 

EST. As noted previously, both of the methods are nonparametric in the interiors of their 

cumulative distribution functions. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The Modified JPM Method (JPM-OS) 

In this section, a method for estimating storm surges via a modified JPM will be developed, 

including estimation of some key climatological characteristics of hurricane tracks, intensities, 

and sizes. The term JPM-OS will be used here to denote this new methodology, since the 

underlying concept of this methodology is to provide a good estimate of the surges in as small a 

number of dimensions as possible, while retaining the effects of additional dimensions by 

including the ε term within the estimated CDF for surges. This approach will also attempt to 

minimize the number of runs required by improving methods used for interpolating between 

combinations of variables in different simulations. 

Estimation of Spatially Varying Probabilities 

In our new approach, the recommended treatment of geographic variation is to use the 

Chouinard et al. (1997) method for determining optimal spatial size for estimating hurricane 

statistics. In this method, the optimal size for spatial sampling is estimated in a manner that 

balances the opposing effects of spatial variability and uncertainties related to sample size. 

Although the final, definitive statistics are still being developed, a brief description of the 

method is included here, along with the preliminary results. 

We begin by estimating the omni-directional statistical properties for storm frequency and 

intensity. Work performed by Gabriel Toro of Risk Engineering showed that the optimal spatial 

sample (kernel) size was in the range of 160 km for frequency analyses, but found that the 

optimal spatial size for intensities reached a plateau above about 200 km and did not drop off 

substantially at higher spatial kernel sizes. For our purposes, we took the basic data set of 22 

hurricanes, which had central pressures less than 955 mb, shown in Figures A3 through A5 and 

defined their locations and intensities along the line shown in Figure A2. Although this line 

includes the west coast of the Florida peninsula for completeness within the analysis, results will 

only be presented for the section of coast west of this peninsula. Our hurricane sample covers the 

interval 1941 through 2005. Appendix C provides a synopsis of some work supporting the 

selection of this period of record. 

For our frequency analysis, we selected a “line-crossing” methodology, rather than an ‘areacrossing” (such as used in the Toro analysis presented in Figure 1) since the frequency of 

landfalling storms is inherently better posed in this context. The location of this line is shown in 

Figure A2 and distance in this system will be referenced in this white paper via an “increment 

number.” This “increment number” is based on integer values of the distance in degrees 

(longitude at 29.5 degrees latitude) as explained in the figure caption in Figure A2. 

After a number of sensitivity studies we were able to show that the results for spatial samples 

for spatial kernels above 250 km or so did not vary markedly and settled on a sample size of 

±3 degrees (333 km) along this line. Results from this analysis are converted into an estimate of 

the frequency of hurricanes (which attain a minimum central pressure of 955 mb or less) making 

landfall within contiguous 1-degree increments along the reference line. Figure 4 gives the 

results of this analysis. As can be seen here, the “line-crossing” frequency estimate is fairly 

consistent with the spatial-area frequency estimates obtained by Toro in Figure 1. It should be 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-30

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

noted that the Toro analysis was based on 52 storms (all storms above Category 2 intensity 

within the Gulf of Mexico) rather than the 22-storm subset used here. 

For each 1-degree increment along the coast, pressure differentials at the time of landfall for 

all storms making landfall within the ±3-degree distance along the reference line were used to 

define a best-fit (conditional) Gumbel distribution, i.e. the distribution of hurricane intensity 

given that a hurricane (with central pressure less than 955 mb) does occur. The Gumbel 

coefficients for the pressure differentials are shown in Figure 5. Combining the storm frequency 

estimates with these values we can estimate the omni-directional probability of intensity along 

the Gulf coast at the time of landfall. 

Figure 6 shows the (smoothed) distribution of the 50-year, 100-year, and 500-year central 

pressures based on the Oceanweather information. Also shown in this Figure is the estimates 

using the same derivation methodology, but based on the official NOAA values for landfalling 

central pressures. As can be seen here, the two methods yield very similar results, except for 

some divergence as the Louisiana-Texas border is approached. This curious aspect of the 

otherwise excellent agreement was found to be related to a single spurious value in the 

Oceanweather files for Hurricane Audrey. Once this was fixed, the region of good agreement 

extended across the entire region. 

Interpretation of these values to “off-coast” values should refer to Figure 3, which suggests 

that the off-coast central pressures should be on the average around 10-15 mb lower that these 

values. For comparison, return periods based on independent analyses performed by David 

Levinson (National Climate Center) suggest that the 100-year central pressure offshore was in 

the range of 894-908 for a broad portion of the central Gulf of Mexico, which is reasonably in 

agreement with the values derived here, based on 1-degree increments of coast. However, in 

performing this comparison, it is essential to bear in mind that the estimates are appropriate for 

the recurrence within a 1-degree section of coastline. To compare the estimates in Figure 6 with 

values for the entire Gulf of Mexico, Peter Vickery performed an independent analysis of 

extremes based on 1) a statistical combination of all of the coastal segments and 2) an extremal 

analysis of the NOAA’s landfalling pressures. As can be seen in Figure 7a, the estimates of 

landfalling central pressures shown in Figure 6 can be used to provide a very consistent estimate 

for landfalling central pressures along the U.S. coast from Texas to the northwest Florida coast. 

And, as can be seen in Figure 7b, a similar analysis for a single 1-degree increment shows, as 

expected, a little more randomness for this (smaller) set of points but a similar good general 

agreement in the tail of the distribution. Figure 7c shows a comparison of another independent 

check on the probability distribution based on a comparison to the results of Toro’s analysis 

performed for FEMA Region 4, along the Mississippi coast. These tests confirm the general 

estimation methodology used here is quite robust for hurricanes along the U.S. Gulf of Mexico 

coasts. 

Several independent analyses over the last several months have shown that storm size is not 

independent of storm intensity; and recently, Shen (2006) has shown that the potential intensity 

achievable by a hurricane is very sensitive to the size of a hurricane eye. Figures 8a and 8b show 

the relationships between Rp and central pressure from Oceanweather analyses of all storms 

exceeding Category 2 within the Gulf of Mexico at their time of maximum strength (52 storms –
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-31

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

shown in Figure 8a) and the 22-storm sample of landfalling storms (Figure 8b). Equation A3 in 

Appendix A gives an estimate of the conditional probability of storm size as a function of central 

pressure. Figure 8 gives the mean angle of storm heading as a function of distance along the 

reference line shown in Figure A2, along with the standard deviation of the heading angles 

around this mean value. The direction convention used here is that a heading of due north 

represents an angle of zero degrees. Storms heading more westerly than due north will have 

positive angles, while storms heading more easterly will have negative angles. These estimates 

were derived by the same spatial averaging procedure used in deriving the central pressures and 

frequencies. A circular normal distribution is used here to represent the storm heading 

probability distribution as a function of location along the reference line. 

Figure 10 presents the estimated forward storm speed as a function of central pressure. This 

figure suggests that storm intensity and the forward speed of the storm are approximately 

independently distributed. However, if we plot forward storm speed as a function of storm 

heading at landfall for the 14 storm subset that intersect with the 29.5-degree latitude portion of 

the reference line in Figure A2 and for the entire 22-storm sample of landfalling storms (shown 

in Figures 11a and 11b), we see that there is a tendency for higher forward speeds to be 

associated with lower storm heading angle (a correlation of 0.52 which is significant at the 0.05 

level of signficance with 21 degrees of freedom in a “Student’s t” test). This is consistent with 

the expected behavior of re-curving storms that become swept up in stronger westerly 

circulations. The primary exception to the overall relationship is Hurricane Betsy, represented by 

the point in the upper right-hand corner of this Figure. This storm moved very fast into the New 

Orleans area after crossing the lower portion of the Florida peninsula. 

Putting all of the pieces of information together, for any point in our five-dimensional 

parameter space (retaining appropriate interrelationships among parameters), we see that the 

final estimates of joint probability densities can be written as 

2

2

12345

0 1 0

1

1

(() )

2()

2

( ()

3

( , , , ,)

[ ( ), ( )] ( ) ( | ) exp exp (Gumbel Distribution) ( )

1 ( |) ( )2

1 ( |) 2

p p

f l

p pfl

p

p

p

R PR

P

p p

v

f l

pc R v x

Fa x a x c ax

pc x

c x ax

pR c e P

pv e

σ

θ

θ

σ π

θ

σ π

Δ − − Δ

−

=Λ ⋅Λ ⋅Λ ⋅Λ ⋅Λ

∂ ∂ ⎧ ⎫ ⎪ ⎪ ⎪⎪ ⎧ ⎫ ⎡ ⎤ − Λ= = = − − ⎨ ⎨ ⎬⎬ ⎢ ⎥ ∂ ∂ ⎪ ⎪ ⎩ ⎭ ⎪ ⎪ ⎩ ⎭ ⎣ ⎦

Λ= =

Δ

Λ= =

2

2

2

2

)

2

( () )

2 ()

4

5

1 ( |) () 2

( )

f

l l

v

x

x

l px e

x

x

σ

θ θ

σ θ

σ π

−

− −

Λ= =

Λ =Φ

where the overbars denote average values of the dependent variable for a specified value of an 

independent variable in a regression equation, 0 1 ax ax ( ) and ( ) are the Gumbel coefficients for 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-32

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

the assumed Gumbel form of the central pressures, and Φ( ) x is the frequency of storms per year 

per specified distance along the coast (taken as one degree in examples presented here). 

Figures 12-14 show three sets of synthesized tracks that are being used in the ongoing New 

Orleans area study. These central tracks (Figure12) essentially mimic the behavior of intense 

landfalling historical storms in the record, while preserving the geographic constraints related to 

land-sea boundaries. These storms preserve the historical pattern of the tracks better than simply 

shifting the same storm tracks east or west along the coast, since they capture the observed 

variations in mean storm angles along the coast. 

Estimation of the ε Term 

Although there may be some degree of nonlinearity in the superposition of tides and storm 

surges, numerical experiments have shown that for the most part linear superposition provides a 

reasonable estimate of the (linearly) combined effects of tides and surges. Thus, the tidal 

component of the ε term, represents the percentage of time occupied by a given tidal stage and 

can be directly derived from available tidal information along the coast. 

Careful analyses appropriate for formulating Holland B parameters for ocean response 

modeling have shown that this parameter falls primarily in the range of 1.1 - 1.6 offshore and 

0.9 - 1.2 at the coast (Appendix E). For Gulf of Mexico hurricanes, a mean value of 1.27 in 

offshore areas is assumed with a standard deviation of 0.15, while at the coast the corresponding 

mean and standard deviation is 1.0 and 0.10, respectively. Via numerical experiments, the 

maximum storm surge generated by a hurricane has been found to vary approximately linearly 

with variations in the Holland B parameter, at least for changes of the Holland B parameter in 

the range of 10-20%. 

Off-coast track variations affect surges at the coast primarily through the effects of these 

track variations on wave fields, rather than by their effects on direct wind-driven surges. As 

noted previously, wave fields tend to integrate wind field inputs over 10’s of hours; 

consequently, off-coast track variations tend to shift the wave fields somewhat while maintaining 

the general form and magnitude of the wave height contours. Near-coast radiation stresses are 

approximately proportional to gradients in wave energy fluxes, which, in turn, can be related to 

the square of the wave height gradient. In shallow water, where contributions of radiation 

stresses to surges are most important, wave heights tend to be depth limited. It is only in the 

incremental region, where larger waves make additional contributions due to increased energy 

losses offshore, that larger wave conditions affect the total wave set-up at the coast. Numerical 

sensitivity studies suggest that once incident waves become much larger than about 10 meters, 

most of the additional energy loss is in depths that do not contribute very much to wave set up. 

For this reason plus the fact that in general the wave set-up term tends to be only about 15-30% 

of the total surge, we expect the effect of storm track variations on wave set-up at the coast to be 

fairly small (due to the fact that surge response is on a much faster scale than wave generation, 

where we noted that the “straight-track” approximation was not very good). We will assume that 

the deviations around the mean surge will be approximately Gaussian. A standard deviation of 

20% of the calculated wave-set up contributions to the total surge (determined by subtracting the 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-33

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

direct wind-only surge from the total surge due to winds and waves combined) will be used 

within this distribution. 

Model errors combined in calibration/verification runs of ADCIRC have shown that this 

combination of model and forcing in the Louisiana-Mississippi coastal area provides relatively 

unbiased results with a standard deviation in the range of 1.75-2.50 feet. Relative errors 

associated with the use of PBL winds increase the value of the standard deviation to 2.00 to 3.50 

feet. This is not to surprising, since the accuracy of HWM’s (the primary measurements to which 

the model results are compare) are quite variable in and of itself. 

Combining all of these terms, under the assumption that they are each independently 

distributed, gives 

1 2 3 4 1 2 3 4 1234 p() ( )( )( )( )( ) ε = +++− δε ε ε ε ε ε ε ε ε ε ε ε ε p p p p dddd ∫∫∫∫

where 

1

2

3

 is the deviation between a storm at a random tide phase and a zero tide level;

 is the deviation created by variation of the Holland B parameter;

 is the deviation created by variations is tracks

ε

ε

ε

4

 approaching the coast; and

ε is the deviation created primarily by errors in models and grids.

Three of the terms 13 4 ε , , and ε ε are treated here as though they are approximately 

independent of the magnitude of the surge, while the remaining term, 2 ε has been found to 

depend essentially linearly with the magnitude of the surge. For a monochromatic tide, the tidal 

elevation distribution, 1 ε , is known to be bimodal distributed around its zero value; however, in 

nature, the effect of combining several tidal components with varying phases is to force the 

distribution toward a unimodal distribution. The probabilities of terms 3 4 ε and ε are assumed to 

be normally distributed; thus, the probability distribution of the sum of these two terms will also 

be a normal distribution with the variance given by the sums of the individual variances of the 

two terms. 

Figure 15 gives a numerical example of the combination of all four terms assuming a storm 

surge of 15 feet, as might be associated with a particular deterministic model execution based on 

a set of track and PBL parameters. As can be seen in this figure, the overall magnitude of these 

effects can add or subtract substantially to the total water depth. In this case, the distribution 

appears similar to a Gaussian distribution, since it is dominated by the term with the largest 

variance (deviations due to the omission of the Holland B parameter); however, the other terms 

have been included within the integral for p( ) ε . Table 1 shows an example of the effect of 

adding this term on expected surge levels for selected return periods. In this example, a Poisson 

frequency of 1/16 was used in combination with a Gumbel distribution, with parameters 

0 1 a a = 9.855 and = 3.63 . For this example, the effect of adding the ε -term is less than 1⁄2 foot 

for return periods up to 175 years and only exceeds1 foot at return periods greater than 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

400 years. However, for risk-based calculations which often include very large return periods 

(1,000-10,000 years), this term can become as large as 2-3 feet, even for the case where the 

effects of all neglected factors are assumed to be distributed around a mean deviation of zero. 

The effect could of course be larger if the deviations were biased. 

From Table 1 and the above discussion, we see that the effect of the ε -term becomes much 

more pronounced at large return periods. Thus, older applications of the JPM that neglected this 

term were probably reasonably accurate at the 100-year return period, but were likely to have 

been progressively biased low at higher return periods. The important points to stress here are 

twofold. First, any neglect or suppression of natural variability in a procedure to estimate 

extremes will lead to some degree of underestimation of the estimated extremes, whether using a 

JPM or an EST approach; therefore, it is important to recognize and attempt to quantify all 

significant factors affecting surge heights at the coast. Second, to avoid making the number of 

dimensions in the JPM unmanageable, the estimated effects of the neglected factors contributing 

to extreme surges should be addressed statistically, such as we have done here via the addition of 

the ε -term to the JPM integral. 

Treatment of Subsidence and Sea-Level Rise 

Rather than treat subsidence and sea-level rise within the ε -term, it is simpler to include this 

in a separate analysis. For purposes of design, as a first approximation to the non-overtopping 

situations, estimates of subsidence and sea-level rise can be added linearly to the expected surge 

levels. Thus, (as a purely hypothetical scenario) if two feet of local subsidence is anticipated 

along with one-foot of sea-level rise over a design lifetime (say 100 years as an example), a 

levee design set for 20 feet would need to “evolve” to 23 feet in order to provide the same level 

of protection at the end of the design lifetime. Other options might be to overbuild the design at 

the outset to account for anticipated subsidence and sea-level rise. In either case, it will be 

critical to constantly monitor the changing water levels to ensure that the design level of 

protection is maintained. 

Sampling of Storm Parameters for the JPM-OS 

In the conventional JPM each simulation was typically treated as representative of its entire 

discrete probability range (i.e. all of the probability for each multi-dimensional box centered on 

its mean position). In these applications, the computational burden was considerably less than 

what is considered appropriate for surge simulations today (see subsequent section on the 

computational effort recommended for today’s applications). Even in the original JPM, however, 

a scaling relationship between the pressure differential of a storm and computed surge levels was 

used to reduce the number of computer runs. This relationship, based on theoretical 

considerations and confirmed numerically in several studies, shows that surges are linearly 

proportional to the pressure differential of a storm at all areas close to the area of maximum 

storm impact. This information can be used effectively to interpolate between two different 

numerical results within the JPM integral. Such an interpolation provides added resolution along 

the pressure differential axis in this integral, which is very important due to the highly nonlinear 

characteristics of the probability of pressure differentials [ ( ) p ΔP ]. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-35

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

In addition to the scaling relationship between surge levels and pressure differentials, the 

JPM-OS attempts to sample the parameter space in a fashion that can be used to estimate surges 

(develop the response surface) in an optimal manner. This method has been developed via 

hundreds of simulations on relatively straight coasts, as well as on coasts with other simple 

geometries, and is in the process of being extended to more complex coasts. It attempts to 

alleviate the need for very closely spaced parameter values in numerical simulations (essentially 

track spacing and number of storm sizes, forward speeds, and track angles considered); thereby 

potentially greatly reducing the total number of computer runs required for JPM execution. The 

initial set of runs for the New Orleans area consists of 152 hurricanes traveling along tracks 

shown in Figures 12-14. To put this number of runs in perspective, since a major storm only 

affects each one-degree section of coast once per 16 years and the section of coast being studied 

is only about 2.5 degrees, this number of hurricanes would only be expected in the simulated 

area every 853 years. Consequently, unless we have selected these storms in a very 

unrepresentative fashion, we expect this number of storms (combined with an accurate 

methodology for surge simulation) to provide a fairly accurate description of the general 

characteristics of hurricane surges at least up to the 500-year return period. A description of the 

parameters of these storms is given in Appendix D, along with a discussion of some scaling 

relationships between storm parameters and surges that have been found in our numerical 

studies. 

Specification of Variations in Pre-landfalling Hurricanes 

Whereas the original JPM considered storm size, intensity, and wind field distribution to be 

constant in storms approaching the coast, the new JPM uses information from recent storms to 

estimate the rate of change of these parameters for pre-landfall conditions. In general these 

trends show that storms tend to fill by about 10-15 millibars (Figure 3), become slightly 

(15-30%) larger (Figure 16) and have less peaked wind speed distributions (Holland B parameter 

decreasing from about 1.27 to around 1.0) over the last 90 nautical miles of coastal water before 

landfall. Since all of our probabilities have been developed based on landfalling characteristics, 

the offshore characteristics must be estimated from a generalized transform 

1 ( , , , ,) ( , , , ,) p f l offshore p f l landfall p PR v x p PR v x J θ θ − Δ =Δ

where J is the Jacobian for the transform from nearshore to offshore conditions. However, since 

1) storm heading during approach to the coast is relatively constant, 2) the forward speeds are 

assumed to be constant during approach to land and 3) the points of intersection (x) are identical 

for each offshore and landfall case, the transform can be viewed in only two dimensions, 

 and ΔP Rp . Details will be given in Appendix D. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-36

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A Computational Methodology for Effective Simulation of Storm 

Surges for Hurricane Inundation Studies 

For completeness, we include here a brief description of the computational methodology that 

has been adopted for calculating inundation levels/probabilities. The first step in this procedure 

is to develop appropriate surge and wave model grids and verify them. The second step is to use 

both “best available” winds and best PBL winds to verify the modeling system performance 

within the area of interest. The “best available” winds, which include all appropriate data 

assimilation and expert analyses, are used to verify the model and grid and to provide calibration 

guidance, if required. The second set of runs with winds from the PBL model are used to 

establish any additional tuning required and to determine the actual modeling error term to be 

used in the ε -term in the JPM integration. Following this procedure, the complete production 

system is exercised for all of the selected combinations of storm tracks and wind parameters 

required for the JPM-OS application. A schematic diagram of this system is shown in Figure 17. 

Step 1 

As seen here, for each defined storm (a track and its time-varying wind field parameters) the 

Oceanweather PBL model (Thompson and Cardone, 1996) is used to construct 15-minute 

snapshots of wind and pressure fields for driving surge and wave models. 

Step 2 

Using “warm-start” condition with all major rivers already “spun up,” initiate ADCIRC 

(version 46.50 or higher as they come online) for simulation of direct wind-driven surge 

component (assuming zero tide). In parallel with the initial ADCIRC runs, execute a largedomain,discrete, time-dependent spectral wave model (WAM or WAVEWATCH) to calculate 

directional wave spectra that serve as boundary conditions for local-domain, near-coast wave 

model runs in Step 3. 

Step 3 

Using initial water levels from ADCIRC, winds that include the effects of sheltering due to 

land boundaries, and spectral boundary conditions from the large-domain wave model, execute 

either STWAVE or SWAN model runs (again using the PBL winds) to produce a wave fields 

and estimated radiation stress fields. 

Step 4 

Using the radiation stress fields from Step 3 added to the PBL-estimated wind stresses, rerun 

the ADCIRC model for the time period during which the radiation stresses potentially make a 

significant contribution to the water levels. In this step, care must be taken to “match” the grids 

in the wave model and ADCIRC model. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-37

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Step 5 

Using the water levels from Step 4, in locations adjacent to structures, a method based on 

Boussinesq modeling (either direct application or interpolations from generic runs) is used to 

provide estimates of the following key information along man-made structures and steep-sided 

natural flood barriers: incremental contribution to the water level at the structure (which should 

be considered as the final water level for inundation efforts at these locations); estimated total 

overtopping at structure; and estimated velocities on the front face, crest, and rear face of the 

structures. The boundary conditions for driving the Boussinesq-based runs are taken from the 

closest points in the nearshore grids used in Step 3. 

Summary 

Some of the key issues addressed in this paper, along with relevant conclusions, are as 

follows: 

1. The Joint Probability Method (JPM) provides a sound method for estimating inundation 

probabilities. However, given the number of degrees of freedom in hurricanes 

characteristics affecting coastal surges and the computational burdens associated with 

coastal surge simulations, it is critical to reduce the number of factors considered to a 

minimum, while maintaining sufficient detail to properly model hurricane wind fields for 

surge prediction. 

2. After a discussion of various alternatives, it is recommended that the same five 

parameters used in older JPM studies (storm intensity, storm size, forward speed of the 

storm, angle of the storm track with the coast, and track location) be used to characterize 

storms for simulating coastal surges. With this number of dimensions within the JPM 

integral, it is essential 1) to allow these characteristics to exhibit observed variations 

during their approach to land and 2) to retain and quantify a statistical “error” term that 

adds the suppressed variability back into the estimated extremes. Previous applications of 

the JPM did not consider this term. 

3. Similar to the Empirical Simulation Technique (EST), it is recommended that uncertainty 

in the stage-frequency relationships be estimated via re-sampling methods. 

4. As discussed in Appendix C and as has been noted in many journal publications, it 

appears that a 40-year cycle is a dominant feature within the recent hurricane record. We 

have experienced one full “high-activity” portion of a cycle and about 2/3 of a second 

“high-activity” portion of a cycle. It is recommended that a comparable proportion of 

“low-activity” years be included within the record being used for estimating inundation 

probabilities. Following this logic, we will use the full “low-activity” interval 

(approximately 30 years) between the two recent “high-activity” intervals plus a period 

equivalent to 2/3 of the recent “low-activity” interval within our sample period. This 

yields 1941 through 2005 as our period of record for coastal inundation analyses. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-38

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

5. The topic of future climatic variability has been dealt with in a recent manuscript by 

Resio and Orelup. In this manuscript, it is shown that the hurricane record within the 

Gulf of Mexico does not exhibit a strong secular trend, in contrast to the record in the 

Atlantic Basin. The Resio and Orelup manuscript showed that the “high-activity” 

intervals dominated the extreme surge population for return periods greater than 50 years 

or so. As a sensitivity study, they investigated the consequences of a doubling of “highactivity” years, even though there is no evidence that such a doubling is imminent. 

Results of this study showed that such a climate scenario would produce about a 12% 

increase in surge levels at the 100-year level, with decreasing effects at longer return 

periods. 

6. Historically, the storms which appear to have had the most impact on coastal areas within 

the Gulf of Mexico have all moved along the central paths shown in Figure 12. Theses 

storms include Rita, Ivan, Camille, Katrina, and Andrew in the historical record from 

1941 through 2005; thus the set of tracks in this figure have been nicknamed the RICKfan. 

7. The JPM-OS represents an attempt to combine statistical information over an interval of 

the coastline in order to gain more confidence in information relevant to the definition of 

extreme surges at a point. This approach also allowed us to incorporate information on 

the general behavior of storms in coastal areas (such as storm decay and variations in 

storm size and the peakedness of the wind distribution along a transect) into our 

simulations of extreme events. 

References, Tables, and Figures for R2007 

ASCE, 1990: Minimum design loads for buildings and other structures, ASCE 7-88, New York. 

ASCE 1996: Minimum design loads for buildings and other structures, ASCE 7-95, New York. 

Borgman, L.E, Miller, M., Butler, L. and R. Reinhard, 1992: Empirical simulation of future hurricane 

storm histories as a tool in engineering and economic analysis, Proc. Fifth Intn’l, Conf. on Civil Engr. 

In the Ocean, ASCE, November, College Station, Texas. 

Cardone, V.J., Pierson. W.J., and E.G. Ward, 1976: Hindcasting the directional spectra of hurricane 

generated waves, J. Petrol. Technol., 28, 385-394. 

Collins, J.I., and M.J.Viehmann, 1971: A simplified model for hurricane wind fields, Paper 1346, 

Offshore Technology Conference, Houston, TX. 

Chouinard, L.E., Liu, C. and C.K. Cooper, 1997: A model for the severity of hurricanes in the Gulf of 

Mexico, J. Waterways, Harbor and Coastal Engr., ASCE, 123, 113-119. 

Chow, S., 1971: A study of the wind field in the planetary boundary layer of a moving tropical cyclone, 

M.S. Thesis, New York University. 

Ho, F.P and V.A. Myers, 1975: Joint Probability Method of Tide Frequency Analysis applied to 

Apalachicola Bay and St. George Sound, Florida, NOAA Tech. Rep. NWS 18, 43p. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-39

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Ho, F.P., Su, J.C., Hanevich, K.L., Smith, R.J. and F.P. Richards, 1987: Hurricane Climatology for the 

Atlantic and Gulf Coasts of the United States, NOAA Tech. Rep. NWS 38, completed under 

agreement EMW-84-E-1589 for FEMA, 194 p. 

Holland, G.J., 1980: An analytic model of the wind and pressure profiles in hurricanes, Mon. Wea. Rev. 

108, 1212-1218. 

Irish, Resio, & Ratcliff, 2006: The influence of hurricane size on coastal storm surges, submitted to J. 

Phys. Oceanogr. 

Kimball, S.K., 2006: A modeling study of hurricane landfalls in a dry environment, Mon. Wea. Rev., 134, 

1901-1918. 

Myers, V.A., 1954: Characteristics of United States hurricanes pertinent to levee design for Lake 

Okeechobee, Florida, Hydromet. Rep. No. 32, U.S. Weather Bureau, Washington, D.C. 

Myers, V.A., 1975: Storm Tide Frequencies on the South Carolina Coast, NOAA Tech. Rep. NWS-16, 79 

p. 

Resio, D.T. and E.A. Orelup, 2006: Potential effects of climatic variability in hurricane characteristics on 

extreme waves and surges in the Gulf of Mexico, submitted to J. Climate. 

Scheffner, N., Borgman, L. and D. Mark, 1993: Empirical simulation technique applications to a tropical 

storm surge frequency analysis fo the coast of Delaware, Proc. Third Intn’l Conf. on Estuarine and 

Coastal Modeling. 

Schwerdt, R.W., Ho, F.P., and R.R. Watkins , 1979: Meteorological criteria for Standard Project 

Hurricane and Probable Maximum Hurricane Windfields, Gulf and East Coasts of the United States, 

Tech. Rep. NOAA-TR-NWS-23, National Oceanic and Atmospheric Administration. 

Shen, W., 2006: Does the size of hurricane eye matter with its intensity?, Geophys. Res. Let., 33, L18813. 

Thompson, E.F. and V.J. Cardone, 1996: Practical modeling of hurricane surface wind fields, ASCE J. 

Waterway, Port, Coastal, and Ocean Engineering, 122, No. 4, 195-205. 

Vickery, P.J., Skerjl, P.F., and L.A. Twisdale, 2000: Simulation of hurricane risk in the U.S. using 

empirical track model, J.Struct. Engr., 1222-1237. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-40

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Table 1. 

Example of Expected Surge Values as a Function of Return Period With and Without 

ε-Term Included 

Return Period (years) Without ε-Term (feet) With ε-Term (feet) 

50 11.98 12.06 

75 13.64 13.90 

100 14.82 15.21 

125 15.74 16.22 

150 16.49 17.04 

175 17.12 17.74 

200 17.67 18.35 

225 18.15 18.88 

250 18.59 19.36 

275 18.98 19.79 

300 19.33 20.18 

325 19.66 20.55 

350 19.97 20.88 

375 20.25 21.20 

400 20.52 21.49 

425 20.76 21.76 

450 21.00 22.02 

475 21.22 22.27 

500 21.43 22.50 
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Figure 1. Analysis of hurricane frequency from Toro (Risk Engineering) from an analysis using an 

optimized spatial kernel. 
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Figure 2. Distribution of the maximum surge heights from ADCIRC simulations for a site in Lake 

Pontchartrain. Storms within the “non-event” asymptote consist of storms which do not make 

landfall close to the site of interest; whereas, storms within the “direct-hit” asymptote represent 

storms that pass very close to the site of interest. The different slopes of these line segments 

suggests that it may not be justifiable to combine these points into an analysis that treats all of 

the points as though they are drawn from a single analysis. 

Direct Hit 

Asymptote

“Non-event” Asymptote 

Katrina
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Figure 3. Central pressure in landfalling storms plotted against distance from the coast. Previously it was 

believed that storm decay began only after landfall. These data from Oceanweather, Inc. show 

clearly that decay begins offshore. 

LAND WATER 
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Figure 4. Frequency of hurricanes along reference line with annotated geographic locators, based on 

22-storm sample. Location along this line can be taken as equivalent to 1-degree increments 

along the coast, with the New Orleans area falling within increment 7. 

New Orleans 

Panama City LA/Texas Border 

Texas/Mexico Border 
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Figure 5. Gumbel coefficients for locations along reference line, based on 22-storm sample. For 

reference, the Gumbel equation is reproduced here in terms of its explicit dependence on x, 

⎧ ⎫ ⎪ ⎪ ⎡ ⎤ Δ − Δ =− − ⎨ ⎬ ⎢ ⎥ ⎪ ⎪ ⎩ ⎭ ⎣ ⎦

0

1

( ) ( | ) exp exp ( )

P ax F Px

a x

, where ΔP is the pressure differential (peripheral 

pressure minus central pressure). It should be recognized that the frequency is assumed to be 

equal to 1 in this equation. Location along this line can be taken as equivalent to 1-degree 

increments along the coast, with the New Orleans area falling within increment 7. 

LA/Texas Border 

New Orleans 

Panama City 

Texas/Mexico Border 
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Figure 6. Distribution of 50-year, 100-year, and 500-year central pressures along the reference line 

shown in Figure A2, using both Oceanweather, Inc. (OWI) data and official NOAA values. 

Location along this line can be taken as equivalent to 1-degree increments along the coast, 

with the New Orleans area falling within increment 7. 
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Figure 7a. Comparison of Vickery’s analysis of the combination of distributions for landfalling central 

pressures from all coastal segments (taken from the NOAA results shown in Figure 6) 

compared to the distribution of all (NOAA) landfalling central pressures within the Gulf of 

Mexico. 

Landfalling Storms Texas to NWFL (inclusive)

860

880

900

920

940

960

980

1000

1020

1 10 100 1000

Reurn Period (Years)

Central Preessure (mbar)

15 Segments Combined as Statistically Independent

All Landfall Storms Since 1941
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Figure 7b. Same as Figure 7a except specific to the 1-degree increment centered on 7. 

Landfalling Storms/One Degree Centered on 7

860

880

900

920

940

960

980

1 10 100 1000

Reurn Period (Years)

Central Preessure (mbar)
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Figure 7c. Independent estimate of storm probabilities in the Mississippi coastal area by Gabriel Toro (for 

FEMA Region 4) compared to estimate based on Gumbel segments developed in this White 

Paper. As can be seen here the mean curve is in very good agreement with Toro’s results. 
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Figure 8a. Relationship between size scaling parameter (Rp) versus Central Pressure for 52 storm set in 

Gulf of Mexico (offcoast; all storms > Cat 2). 
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Figure 8b. Relationship between size scaling parameter (Rp) at landfall versus Central Pressure for 22 

storm set in Gulf of Mexico (offcoast; all storms with central pressure < 955 at time of minimum 

pressure in the Gulf of Mexico). 
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Figure 9. Plot of mean storm heading angle and standard deviation around this angle as a function of 

location along reference line. Distance along the x-axis can be taken as equivalent to 1-degree 

increments along the coast, with the New Orleans area falling within increment 7. 
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Figure 10. Plot of forward speed of storm at landfall versus central pressure at landfall. 
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Figure 11a. Plot of storm heading and forward speed at time of landfall for only central Gulf landfalling 

storms. 
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Figure 11b. Plot of storm heading and forward speed at time of landfall for the entire 22-storm sample. 
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Figure 12. Central angle tracks (RICK-fan set). 
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Figure 13. Tracks from southeast at 45-degree angle to RICK-fan set. 
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Figure 14. Tracks from southwest at 45-degree angle to RICK-fan set. 
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Figure 15. Percentage of deviations per 0.1-foot class as a function of deviation in feet. 
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Figure 16. Increase in hurricane size during approach to coast, as seen in recent, well-documented 

storms. 
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Figure 17. Schematic diagram of “production” modeling system for coastal inundation applications. 

Probability Characteristics for Set of PBL Parameters and Tracks

(including potential climate variability) 

Wind Field – Wind Stresses

Surge Models Wave Models

Water Levels, Overtopping, Loads on Structures, etc. 

Winds: 

Oceanweather PBL

Offcoast Waves: 

 WAM/Wavewatch 

Nearshore Waves: 

STWAVE/SWAN 

Local Scale Waves: 

Boussinesq - Parametric

Unified Grid 

Surge model: 

ADCIRC with 

wave set-up 

(version 46.50) 

Coupling

Integration of JPM Probabilities for Inundation Threat 
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APPENDIX A of R2007 

Expedient Estimation of Return Periods for Specific Historical 

Storms 

On one hand, the most accurate characterization of the return period for a specific storm 

water level is to wait until the stage-frequency relationships have been developed and see where 

the water levels from that storm fall on this curve. Inspection of such an approach would show 

that water levels from a specific storm would fall at different return periods at different locations. 

A different, albeit more approximate approach which does not have to wait until the final results 

of simulations are complete, is to treat the storm water level potential in terms of the primary 

parameters affecting peak surge levels, typically storm intensity, storm size, angle of approach to 

the coast, forward storm speed, the geometry of the coast, and the offshore bathymetric slope. 

Application to Hurricane Katrina 

Figure A1, taken from Irish et al. (submitted for journal publication), shows the dependence 

of peak surge on storm intensity (peripheral pressure minus central pressure) and size (scaling 

radius for the pressure field). These values were generated from ADCIRC simulations using the 

Oceanweather Planetary Boundary Layer (PBL) model for wind fields. In these runs, storms 

were held at constant intensity and size during the approach to the coast. This figure is 

appropriate for a very shallow offshore slope (1:10000), found to be approximately 

representative of the continental shelf east of the Mississippi River in this area. Sensitivity 

studies of the effects of storm speed and storm approach angle showed that these factors were of 

secondary influence on surges at the coast, at least within the ranges of expected values for large, 

intense storms. Sensitivity studies also showed that although the coastal geometry in the New 

Orleans area (the presence of the Mississippi River Delta and river levees) tends to modify local 

storm surge levels, this factor did not change the relative ranking of the different storms. 

Assuming that we can neglect the secondary factors, the relative ranking, and thus the return 

period for a specific event, can be deduced from a combination of Figure A1 and a specification 

of the joint probability structure of storm intensity and storm size. The estimation of the latter is 

a key part of the same methodology being used in the JPM-OS application in this area. 

The frequency of storms along a coast has units of number of storms per year per length of 

coastline. For our application, we will assume that a storm within 1/2 degree longitude is the 

relevant parameter for our purpose here. This is equivalent to assuming that the surge within 

±1/2 degree remains sufficiently close to the peak value that is can be considered approximately 

equal to the peak (within 10%). Although a continuous method can be used for this estimation 

method, the results were quite similar and the discretized method is easier to explain. Figure A2 

shows a line along the Gulf of Mexico coastline that will be taken as the “sample line” for 

landfalling conditions in our analysis. A centered, running average of landfalling conditions over 

a distance of 6 degrees along this line was used to estimate the frequency of storms and to 

accumulate samples for extremal analysis within contiguous 1-degree (longitude – at 

29.5 degrees) increments along the coast. The New Orleans area is located in Increment 7 along 

the coast in this scheme. 
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The storm sample used in this analysis is the 22-storm sample shown in Figures A3-A5 

(essentially all storms with central pressures less than 955 mb during their transit through the 

Gulf of Mexico since 1941). Eleven of these storms fell within ±3 degrees of the center of the 

New Orleans section of coast. Since this is not a very large number, we restrict our analysis to a 

2-parameter (Gumbel) distribution here. The conditional Gumbel distribution of hurricane 

pressure differential was found to be given by 

A1. (| ) z e F p hurricane e − − Δ =

where F is the condition cumulative distribution function (i.e. the expected CDF given a 

hurricane) and z is given by the best 2-parameter fit 

0

1

p a

z

a

Δ − =

where a0 and a1 are the distribution parameters and Δp is the pressure differential. For 

Increment 7 along the coast, the value of a0 and a1 are 56.557 and 16.463, respectively. The 

frequency of occurrence of hurricanes for this coastal segment is 0.0486 per year per degree (or 

once per every 20.6 years, or so). 

Since, the contours in Figure A1 denote lines of equal surge, The probability of a storm 

capable of generating a storm surge equal to or greater than that produced by Hurricane Katrina 

can be estimated by the sum of the probabilities in regions A, B, and C, delineated by different 

Rp limits in Figure A6 or 

F ABC ( ) η =++ ∫∫∫

It can be shown that all three of these integrals can be written in a common form 

A2. 

2

1 ( )

( ) (,, ) ( | )

R

p p

R pR

F p A B C P R p d pdR

p

∞

Δ

∂ Δ = ΔΔ

∂Δ ∫ ∫∫

where the limits of the first integral are made to match the region shown in Figure A6. The 

conditional probability for size is given by a Gaussian distribution of the form 
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A3. 

2

2 1 (|) ( )2

( )

( )

x

p

p p

PR p e

p

where

RRp

x

p

σ π

σ

−

Δ =

Δ

⎛ ⎞ − Δ = ⎜ ⎟ Δ ⎝ ⎠

In this equation a linear regression ( R

p =14. 0.3* 110. + −Δ ( p) – with units for Rp and Rp

 in 

nautical miles and units for Δp in millibars implied) was used to represent the conditional mean 

for storm size and the standard deviation was taken as ( ) 0.44 ( ) p σ Δp = Δ R p . The sum of all three 

integrals is 0.0518, which can be interpreted as slightly more than 1 hurricane in 20 can produce 

a surge of Katrina’s level in the New Orleans area. When 1 is divided by the product of the two 

frequency parameters, it yields an estimate of 397.6 years for the return period of a storm 

capable of producing a surge of comparable magnitude to Katrina. For practical purposes, 

Katrina would seem to fall in the range of a 400-year storm, in terms of storm surge generation. 

It should be noted that the methodology described here pertains to the area of maximum surges 

within a storm; consequently, the surges at some distance from area of maximum surge will 

certainly not be 400-year surges for those points. 

Application to Hurricane Rita 

The same procedure outlined above, modified to consider the different offshore slope, was 

applied to Hurricane Rita, with a landfall in Increment 10 of our analysis. A slope of 1:1000, 

which is characteristic of the shelf region in the area west of the Mississippi River, was adopted 

here in the idealized surge modeling referenced here. The return period for this storm from this 

analysis is 89.7 years. Again, for practical purposes, this can be taken to imply that Rita was 

somewhere in the neighborhood of a 90-year storm, in terms of storm surge. 
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Figure A1. Contour plot of peak surge levels along a straight coast with a constant offshore slope as a 

function of storm size (Rmax = Rp in our terminology) and storm intensity (peripheral pressure 

minus central pressure). The storm values used in this plot are the offshore conditions, rather 

than the conditions at the coast. 
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Figure A2. Location of line for analysis of hurricane landfalling characteristics. Throughout this white 

paper, one-degree increments of distance along this line from east to west, with the “zerovalue” taken at -83 degrees longitude, will be used as a locator for discretized sections of 

coast. In this convention, the increment number for any section being analyzed is given by 

Nincrement = Integer (-83-longitude). For example, any point with a longitude less or equal to -83 

and greater than -84 would fall in increment 0, any point with a longitude less than or equal to 

-85 and greater than -85 would fall in increment 1, etc. 
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Figure A3. Tracks of all hurricanes (1941-2005) making landfall in the Western Gulf of Mexico for storms 

that attained a central pressure of 955 millibars or lower during its transit through the Gulf of 

Mexico. 
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Figure A4. Tracks of all hurricanes (1941-2005) making landfall in the Central Gulf of Mexico for storms 

that attained a central pressure of 955 millibars or lower during its transit through the Gulf of 

Mexico. 
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Figure A5. Tracks of all hurricanes (1941-2005) making landfall in the Eastern Gulf of Mexico for storms 

that attained a central pressure of 955 millibars or lower during its transit through the Gulf of 

Mexico. 
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Figure A6. Regions of size-intensity domain expected to contribute to surges greater than or equal to that 

of Hurricane Katrina. The Blue “star” represents Katrina’s characteristics at time of landfall, 

with a somewhat lower intensity than shown in Figure A1. 

A

B

C
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APPENDIX B of R2007 

Selection of Wind Model for Coastal Surge Estimation 

Ocean response models for waves and surges require winds at a constant reference level with 

some suitable averaging interval (typically around 20-30 minutes) and a specified drag law to 

convert these winds into estimates of momentum fluxes from the atmosphere to the ocean. Two 

different models (Thompson and Cardone, 1996; Vickery et al., 2000) were investigated for 

application to coastal surge modeling as part of a general investigation of modeling for coastal 

inundation. 

The dynamic (Planetary Boundary Layer – PBL) model of Thompson and Cardone (1996) 

has been selected for use in estimating hurricane wind fields. In older storms, the “best-estimate” 

wind fields will be virtually identical to the initial guess wind fields due to the lack of 

information available for assimilation into the PBL wind fields. Comparisons of Oceanweather 

and Vickery PBL winds from Hurricane Katrina, along east-west and north-south transects 

through the center of the storm, to Oceanweather’s “best-estimate” winds are shown in Figures 

B1 and B2, respectively. Figure B3 shows a similar set of comparisons for both the east-west and 

north-south transects for Hurricane Betsy. As can be seen in these figures, Oceanweather’s PBL 

winds capture most of the broad-scale structure of the wind fields. 

The Vickery PBL model does not perform as well as the Oceanweather PBL model in 

comparisons with the Oceanweather “best-estimate” winds. This is not too surprising, since the 

Oceanweather analysis “best-estimate” analysis is likely to have been considered in deriving the 

PBL parameters. The abrupt drop-off of wind speeds in the Vickery wind is an artifact of the 

version of the Vickery code that was available for testing during the time of this comparison. 

This problem is being remedied and is not a property of the general solution capabilities inherent 

to the Vickery model. 

Since the relaxation time for coastal surge is on the order of hours rather than 10’s of hours, 

as for the generation of waves, offshore complexities of storm tracks tend to affect wave fields 

much more than they affect coastal surges. Since tracks of major storms tend to lack the 

complexity exhibited by minimal hurricanes and tropical storms and since wave generation tends 

to produce a wave field that represents the integrated effects of winds over many hours, the 

effect of track variation on wave fields should be rather minimal. It should also be recognized 

that PBL winds have long been shown to be capable of providing accurate wind field estimates 

for purposes of hindcasting waves in hurricanes (Cardone et al., 1976). 

The use of PBL winds is also consistent with the approach of Vickery et al. (2000) or 

estimating wind hazards in U.S. coastal areas, which were found to provide very reasonable 

wind estimates when compared to both onland and offshore wind observations. Given all of the 

problems associated with using historical wind fields with indefinite degrees of freedom in their 

formulation, it would seem that a very careful study of the differences between PBL-driven 

surges and “best-estimate”-driven surges should be undertaken for a small set of hurricanes 

before any clear advantage to using “best-estimate” wind fields for surge prediction can be 

claimed. 
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A very important improvement pertinent to PBL wind field estimation is the availability of 

carefully estimated size parameters for all historical storms (from Oceanweather). These 

parameters were compiled specifically for driving ocean response models (e.g. wave and surge 

models) and allow improved estimation of the joint probability of intensity-size relationships in 

Gulf of Mexico hurricanes. 

As a final element of this Appendix, the Holland B parameter will be briefly described. This 

description is excerpted from the section on Meteorology in the Coastal Engineering Manual

(Resio, Thompson, and Bratos), which can be found online for reference. 

Myers (1954), Collins and Viehmann (1971), Schwerdt et al. (1979), and Holland (1980) all 

present descriptions and justifications of various parametric approaches to wind-field 

specification in tropical storms. Cardone et al. (1992) use a modified form of Chow's (1971) 

moving vortex model to specify winds with a gridded numerical model. However, since this 

numerical solution is driven only by a small set of parameters and assumes steady-state 

conditions, it produces results which are of similar form to those of parametric models. The 

Holland model differs from previous parametric models in that it contains a parameter (the 

Holland B parameter) which allowed the peakedness of winds in a hurricane to vary. This model 

will be described here to demonstrate the role of the Holland B parameter in this model. 

In the Holland model, hurricane pressure profiles are normalized via the relationship 

B1. c

n c

p - p

 = 

p - p

β

where 

 p is the surface pressure at an arbitrary radius (r); 

 pc is the (surface) central pressure in the storm; and 

 pn is the ambient surface pressure at the periphery of the storm. 

Holland showed that the family of β-curves for a number of storms resembled a family of 

rectangular hyperbolas and could be represented as 

B2. 

1

1

ln( )

 or

exp

 or

exp

B

p

B

p

B

p

r A

A

r

A

r

β

β

β

−

−

−

=

⎛ ⎞ = ⎜ ⎟ ⎜ ⎟ ⎝ ⎠

⎛ ⎞ − = ⎜ ⎟ ⎜ ⎟ ⎝ ⎠
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-73

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

where 

 A is a scaling parameter with units of length; and 

 B is a dimensionless parameter which controls the peakedness of the wind speed distribution. 

This leads to a representation for the pressure profile as 

B3. c nc ( )exp B

p

A

pp p p

r

⎛ ⎞ − =+ − ⎜ ⎟ ⎜ ⎟ ⎝ ⎠

which then leads to a gradient wind approximation of the form 

B4. 

1/2

2 2 ( )exp

4 2

n c B

p

gr B

a

A AB p p

r r f rf U

ρ r

⎡ ⎤ ⎛ ⎞ − ⎢ ⎥ − ⎜ ⎟ ⎜ ⎟ ⎝ ⎠ = +−

⎣ ⎦

where 

Ugr represents the gradient approximation to the wind speed. 

In the intense portion of the storm, this equation reduces to a cyclostrophic approximation (a 

flow in which the pressure gradient force is balanced only by centrifugal acceleration) given by 

B5. 

1/2

( )exp n c B

p

c B

a

A AB p p

r

U

ρ r

⎡ ⎤ ⎛ ⎞ − ⎢ ⎥ − ⎜ ⎟ ⎜ ⎟ ⎝ ⎠ =

⎣ ⎦

where 

Uc represents the cyclostrophic approximation to the wind speed; which yields explicit 

forms for the radius to maximum winds as 

B6. 

1

max B R = A

where Rmax is the distance from the center of the storm circulation to the location of maximum 

wind speed, compared to pr which is the pressure-scaling radius. 

The maximum wind speed can then be approximated as 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-74

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

B7. 

1/2

1/2

max ( ) n c

a

B U pp

ρ e

⎛ ⎞ = − ⎜ ⎟ ⎝ ⎠

where 

Umax is the maximum velocity in the storm; and 

 e is the base of natural logarithms, 2.718. 

If B is equal to 1 in this model, the pressure profile and wind characteristics become similar 

to results of Myers (1954), Collins and Viehmann (1971), Schwerdt et al. (1979), and Cardone 

et al. (1976). In the case of the Cardone et al. model this similarity would exist only for the case 

of a storm with no significant background pressure gradient. Although the Cardone et al. PBL 

model initially did not consider the effects of the Holland B parameter, it now does include this 

term in its formulation (Thompson and Cardone, 1996). The Vickery et al. (2000) model also 

includes the Holland B term in its formulation. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-75

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure B1. Plot of wind speeds (in meters per second) along east-west transects through Hurricane 

Katrina. HBL denotes wind speeds from the Vickery et al. (2000) hurricane boundary layer 

model. PBL denotes wind speeds from the Thompson and Cardone (1996) planetary boundary 

layer (PBL) model; and OWI denotes wind speeds from the “best-available” wind speeds from 

analysts at Oceanweather, Inc., which include the HWIND inputs from Mark Powell in NOAA’s 

Hurricane Research Division. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-76

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure B2. Plot of wind speeds (in meters per second) along north-south transects through Hurricane 

Katrina. HBL denotes wind speeds from the Vickery et al. (2000) hurricane boundary layer 

model. PBL denotes wind speeds from the Thompson and Cardone (1996) planetary boundary 

layer (PBL) model; and OWI denotes wind speeds from the “best-available” wind speeds from 

analysts at Oceanweather, Inc., which include the HWIND inputs from Mark Powell in NOAA’s 

Hurricane Research Division. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-77

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure B3. Plot of wind speeds (in meters per second) along north-south (N-S) and east-west (E-W) 

transects through Hurricane Betsy. Vick denotes wind speeds from the Vickery et al. (2000) 

hurricane boundary layer model. OWI denotes wind speeds from the Thompson and Cardone 

(1996) planetary boundary layer (PBL) model; and “best-available” denotes wind speeds from 

the “best-available” wind speeds from analysts at Oceanweather, Inc. 

V2

V3

18 20 22 24 26 28 30

5

10

15

20

25

30

35

40

vick

owi

best

Betsy 0912 N-S

Frame 002 ⏐ 06 Sep 2006 ⏐

V1

V3

-95 -90 -85 -80

5

10

15

20

25

30

35

vick

owi

best

Betsy 0912 EW

Frame 001 ⏐ 06 Sep 2006 ⏐
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APPENDIX C of R2007 

Selection of Period of Record for Surge Estimation and 

Consideration of the Effects of Climatic Variability on Surge 

Extremes 

An analysis of climatic variability was undertaken by Resio and Orelup (submitted for 

publication to J. Climate). Enclosed here are some of the findings relevant to the selection of the 

period of record and the estimated effects of climatic variability on expected surge extremes. 

Three fundamental data sets are used in this study: 1) sea-level pressures (SLP’s) from the 

NOATL-tropic data set (a sub-domain of the total NCEP SLP data set that covers from 0o

 to 40o

N latitude and from 5o

 to 110o W longitude); 2) sea surface temperature (SST) data downloaded 

from Extended Reconstructed Sea Surface Temperature (ERSST); and 3) information on 

hurricane characteristics taken from Oceanweather, Inc files, now available in the public domain. 

Details on these data sets are available on appropriate web sites. 

Empirical Orthogonal Functions (EOF’s) have long been recognized as a powerful tool for 

encapsulating natural patterns within the atmosphere. In this study we used data from the 

1950-2005 period (56 years) and defined five-day mean sea level pressure (SLP) fields on a 2.5o

by 2.5o

 grid. This resulted in 73 five-day intervals for every year without a leap year. Leap year 

was handled by adding that day into the time interval starting on February 25th, which created 

one element encompassing six days once in every four years. Given that we were not interested 

in the seasonality of hurricane but rather in inter-annual and longer variability, we defined mean 

pressure fields for each five-day interval throughout the year, with the average taken over the 

56 years included in this analysis. Calculated mean pressure fields for each five-day interval 

were subtracted from individual mean five-day pressure fields to produce a set of 73 x 56 

pressure fields that were input into an EOF analysis. 

The SST data used here represent a subset of the total ERSST data set and covers from 18o

 to 

30o

 N latitude and from 58o to 98o W longitude on a 2o by 2o grid, with a land mask that restricts 

the data to only water points. Although SST patterns within the Gulf exhibit considerable spatial 

variability, it is not clear that the variations in the spatial characteristics of these patterns play a 

major role in the inter-annual variability of hurricane genesis and/or development. Consequently 

such variations are not considered here. Instead, mean monthly data for the entire Gulf of 

Mexico region for July through October were averaged together to provide a single measure of 

sea surface temperature for each hurricane season from 1950 through 2005. 

The data set for hurricane characteristics includes estimates of six-hourly storm position, 

along with several parameters that relate to hurricane shape, size and intensity. Unlike previous 

data sets which have focused on short-duration (typically 1-minute maximum) wind speeds from 

flight level, this new data set also contains estimates of the highest sustained (30-minute 

average) surface-level (10-meter) wind speeds along the path of the storms. Since these are the 

appropriate winds for driving ocean response models, they provide a much more direct measure 

of hurricane surge and wave production. In the earlier storms in this data set (prior to 1990) these 

wind estimates were derived primarily from simulated wind fields based on a “slab model” of the 

lowest region of the atmosphere combined with a planetary boundary layer model (Thompson 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-79

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

and Cardone, 1996). For most of the later storms these wind fields have been extensively 

reworked by analysts to assimilate available measurements. 

Most past studies of climatic variability have used storm frequency (sometimes stratified by 

Saffir-Simpson scale) to categorize storm activity in each year. However, for our purposes we 

will define a single parameter that incorporates both intensity and frequency into one measure of 

hurricane activity. This measure of annual hurricane activity is obtained by calculating the 

estimated kinetic energy for each storm passing through the Gulf of Mexico at the time of its 

maximum intensity and then adding all maxima within a given year. At a fixed time, the total 

kinetic energy in a hurricane can be related to storm size and storm intensity as 

C1. 

2 2

max 1 2 0

max

2 2

max max

~ (, ) (, )

given that ( , ) ( ) then

~

k

k

E V x y dxdy V r rd dr

r Vr V

R

E VR

θ θ

θ φ φθ θ

=

⎛ ⎞ = − ⎜ ⎟ ⎝ ⎠

∫∫ ∫∫

where Vmax is the maximum (30-minute average, 10-m) wind speed within the storm and Rmax is 

the radius to maximum winds at the same time. The value of the parameter Ek at the time of 

maximum wind speed during a storm’s passage through the Gulf of Mexico provides a good 

integrated measure of the storm intensity and size at the time of the storm’s maximum intensity. 

Summing all values of Ek for a season yields a surrogate for combined number, size, and 

intensity of storms in a year. 

Figure C1 shows the cumulative hurricane kinetic energy per season as defined previously in 

this paper, smoothed over a running five-year period. Since we are trying to extend the data as 

long as possible, the “smoothed” data at either end of this record is defined only in terms of the 

existing data within the five-year window. For example, the 2005 data considers only data from 

2003, 2004, and 2005 in its average. Thus as the ends are approached, a slight bias is created in 

terms of the mean position of the years contributing to this mean, culminating in a one-year 

displacement at the beginning and end of the analysis along with a reduced averaging window. 

This Figure shows two very notable peaks, one that commenced in the late-1950’s and persisted 

until about 1970 and a second that began around 2000 and has persisted through 2005, with a 

broad trough in Gulf hurricane activity between these two peaks. There does not appear any 

strong secular signal within this record. The first of these periods coincides with the very active 

hurricane seasons that included Hurricanes Carla, Hilda, Betsy, Beulah and Camille that 

devastated much of the Gulf coast during the 1960’s; while the second contains the recent set of 

intense hurricanes, including Lili, Charley, Ivan, Dennis, Katrina, and Rita. 

Figure C1 shows that there are two intervals of high hurricane activity separated by an 

interval of relatively low hurricane activity in the years from 1950 to 2005. If this cycle is real, it 

seems advisable to use a sampling scheme that preserves the expected long-term ratio of highactivity and low-activity years. Since we obviously want to retain both the early (1960-1970) 

interval of high activity latest (partial?) interval (2000-2005) of high activity in our sample, we 

should include about 2/3 of the 30-year low-activity interval preceding 1960 in order to achieve 
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this balance. This suggests that including approximately 20 years of the 30-year low-activity 

interval (1941-1960) is appropriate. Thus, the total record sample length recommended for surge 

estimation in the Gulf of Mexico is 1941-2005. 

Figure C2 shows estimated return periods for central pressures for the total sample within the 

Gulf of Mexico for different estimation methods. The estimation methods examined are 1) 

analysis of only quiescent years (Group1); 2) analysis of only active years (Group 2); 3) analysis 

of both quiescent-year and active-year samples, treated as though they are drawn from separate 

populations (Equation 4 from the Resio and Orelup manuscript, reproduced below); and 4) 

analysis of both quiescent-year and active-year samples as though they are a single population. 

Eq. 4 from Resio and Orlup: 

C2. 

n

n

1 ( ) 1 ()

where

( ) is the return period for x;

 is the frequency of storms in Group n;

F ( ) is the cumulative distribution function (CDF)

 for storms in Group n; and

 is the pro

nnn

n

T x

F x

T x

x

λ β

λ

β

= −∑

portion of years in Group n.

These results combined with integrations of the full probability integral for surges along a 

straight coast from the Irish et al. computer runs yields the results shown in Table C1 for the 

expected variations in wave and surge extremes within the Gulf of Mexico, given a doubling of 

the high-activity years (from 1⁄4 of the total time to 1⁄2 of the total time). There is nothing in the 

Gulf of Mexico record that suggests such a scenario is imminent, so this prediction should 

probably be taken as an upper limit of what could happen rather than what will happen. 
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Figure C1. Plot of estimated cumulative kinetic energy for all storms at time of maximum surface winds 

within each year: 1941-2005. The units of ε are l

4

/t

2

, since we have factored out the mass term 

(consistent with equation C1). Note: ε is termed Ek in Appendix C. 
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Figure C2. Estimated return periods for 4 separate analyses: Low-Activity years only (Group 1 – red line); 

High-Activity Years only (Group 2 – green line); All years into a single analysis (light blue line); 

and estimate based on combined analysis of two populations (equation 4 from Resio and 

Orelup - dark blue line). 

Table C1. 

Estimated changes in extreme waves heights and surges for selected return periods, 

given a doubling of years with high hurricane activity. 

Return Period (years) Change in Wave Height (percent) Change in surge (percent) 

25 +15 +18 

50 +13 +16 

100 +12 +15 

250 +11 +12 

500 +10 +9 
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APPENDIX D of R2007 

Selected Storm Sample for Simulations 

In this Appendix, the terms Rmax (storm-size scaling radius) and Cp (central pressure) should 

be taken as equal to Rp and Cp, as defined elsewhere in this white paper. Tracks 1-5 are defined 

as shown in Figure 12 of the main text. This set of tracks has been nicknamed the “RICK-fan” 

since it mimics the tracks of Rita, Ivan, Camille, and Katrina. Track 1 is the westernmost and 

Track 5 is the easternmost track in this set. Tracks 1b-4b fall midway between the five primary 

tracks. One primed track 4’ is located about 40 nm north of the landfall point for Track 4 to 

allow for storms entering the N-S aligned portion of the coast in this area. 

One of the issues affecting our storm selection is implicit in the spacing of the tracks shown 

in Figure 12. This spacing is approximately 0.6 degrees longitude at a latitude of 29.5 degrees, 

equivalent to a distance of about 31 nautical miles in the along-coast direction. Studies of surge 

response on idealized, open coasts have shown the distribution of surges along the coast scale 

quite nicely with this parameter for a wide range of storm sizes and offshore slopes (Figures D1-

D3). Figure D4 provides a plot with all three of the storm sizes plotted together. A second factor 

which influences along-coast surge variations is the presence of large geometric features in the 

coastal configuration, such as the Mississippi River, its deltas, and the river levees south of New 

Orleans. Figure D5-A shows a general set-up for a numerical study to examine this effect. 

Figure D5-B shows the distribution of surges along the coast. It is apparent in this figure that the 

effect of the land protrusion is to add a second scale to the along-coast surge distribution. In the 

New Orleans area, this scale is likely to be of more significance than the simple open coast 

scaling. In the New Orleans area, this effect will tend to dominate the second source of along 

coast variation for storm within about 30 nm or so of the eastern edge of the land protrusion. In 

this context, the storm track spacing for the New Orleans area should be quite reasonable. 

In the more general case of storms along a straight coast, any single storm will produce a 

patter with a shape similar to those shown in Figures D1-D3. Figure D6 shows a plot of where 

surges from two storm tracks might fall if they were separated by 31 nautical miles on each of 

these figures. This represents the worst case in which neither of the simulated storms captures 

the peak condition that could be generated by these hurricanes. This case is the worst possible 

since it has the maximum distance from the peak on both sides. Since tracks are assumed to be 

distributed uniformly along the coast, the expected value would be the mean of all points 

between the two lines. The bias introduced into the estimated extreme from this set of tracks can 

be estimated as 

D1. 2

1

1 2

1 2

1 2

2 1

( ) 2

( ) 1 ( ) ( )

x

x

E

E

x dx

x x

η η

η η

η

η −

+

+ =

− ∫

which can be calculated numerically for each of the different radii to maximum winds 

considered. For the three cases shown, the calculated values are 0.794, 0.907, and 0.959 for the 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-84

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10 nm, 20 nm, and 30 nm cases, respectively, which translates into low biases of approximately 

20%, 9% and 4% in these “worst-case” situations. It should be noted that the sampling pattern 

itself should not introduce a bias in either direction (high or low); consequently, in other sections 

of the coast, where the exact peak is attained, the simulated value is actually higher than the 

denominator in this equation. Historically, this problem of low and high values along the coast 

has been addressed by some sort of smoothing along the coast. The smoothing interval of 

approximately max

2

R should work relatively well for this purpose. 

Along each of the tracks modeled, the central pressure is allowed to vary during a simulated 

intensification interval until its intensity reaches a plateau. This plateau is maintained until the 

storm comes within 90 nautical miles of the coast at that time, the pressures decay according to 

the (linear interpolation) relationship 

D2. 

00 0

0

0

( ) ( ) (1 )

where

 is the central pressure at s 

 is the distance along the storm track, with located 90 nm from landfall

 is an interpolation multiplier (=1 at 90 nm from landfal

p p decay

p

Cs Cs P

C

s s

λ λ

λ

= −− Δ

l and =0 at landfall)

 is the total change in central pressure over 90 nm approach to landfall ΔPdecay

The pressure decay term is somewhat dependent on storm size, so the following relationship 

was used to represent this term 

D3. 6 (with given in nautical miles)

constrained by ( ,18); ( ,5)

decay p p

decay decay decay

PR R

P Max P Min P

Δ =− ′

Δ= Δ Δ ′ ′

Once a storm is one hour past landfall the pressure decay factor due to Vickery is applied 

D4. 

0

0

where

where

 is the local pressure differential 

 is the pressure differential one hour after landfall

 is an empirical constant

 is time after landfall minus 1 hour

p

a t

CP P

P Pe

P

P

a

t

δ

δ δ

δ

δ

∞

− Δ

= −

=

Δ
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-85

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As noted in the main section of this white paper, Rmax and the Holland B parameter are 

allowed to vary linearly over the same distance as Cp for all storms except the smallest storm 

class used in this application. For that class (Rmax = 6nm), the storm is assumed to retain its 

intensity, its size, and its Holland B parameter all the way to landfall. Figure D6 shows a typical 

variation of storm parameters for a storm in which the characteristics vary during their approach 

to the coast. This figure shows the characteristic “spin-up” time for the hurricane (based on 

historical times required to reach peak storm intensities) and the variations during its approach to 

the coast. 

Four different Cp – Rp (central pressure – size scaling radius) combinations are defined as 

A. 3 Cp and 3 Rmax values 

Cp=960 Rmax = (21.0, 35.6, 11.0) 

Cp=930 Rmax = (17.7, 25.8, 8.0) 

Cp=900 Rmax = (14.9, 21.8, 6.0) 

The loop structure for JPM run sequencing is (from inner loop working outward) 

- Rmax 

- Cp 

- Track 

Thus, the sequence of runs for any track using Combination A is 

 Cp Rmax 

1. 960 15.4 

2. 960 21.0 

3. 960 35.6 

4. 930 11.7 

5. 930 17.7 

6. 930 25.8 

7. 900 6.0 

8. 900 14.9 

9. 900 21.8 

B. 2 Cp and 2 Rmax values 

Cp=960 Rmax=(18.2,24.4) 

Cp=900 Rmax=(12.5,18.4) 

The sequence of runs for any track using Combination B will be 

 Cp Rmax 

1. 960 18.2 

2. 960 24.4 

3. 900 12.5 

4. 900 18.4 
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C. 2 Cp and 1 Rmax values 

Cp=960 Rmax=17.7 

Cp=900 Rmax=17.7 

The sequence of runs for any track using Combination C will be 

 Cp Rmax 

1. 960 17.7 

2. 900 17.7 

D. 1 Cp and 1 Rmax value 

Cp=930 Rmax=17.7 

Defining three angles as the central angles in the RICK-fan +/- 45 degrees should cover most 

of the important range for estimating the response surface of the surges. With the secondary 

variables (tidal phases, Holland B variations, wind field variations around the PBL central 

estimate, etc.) added to the integral, this should provide a very reasonable estimate of the surge 

CDF. Figures 13 and 14 in the main text provide the geographic information on these tracks. The 

tracks approaching the New Orleans area from the southeast are extremely similar to the tracks 

of the 1947 Hurricane, Betsy, and Andrew in this area. During the 1941-2005 interval, no tracks 

approached New Orleans from the southwest; however, other storms such as the 1893 storm did 

approach New Orleans from this direction. In fact, the 1893 track is fairly similar to one of the 

hypothetical tracks out of the southwest. A track from this direction represents the fact that these 

storms have to become caught up in the more westerly flow (winds blowing toward the east). For 

a storm to maintain its strength it cannot move too far west or too close to land; consequently, 

the track of a major storm is constrained somewhat to come from the region from which all the 

hypothetical (+45 degree) tracks emerge in order for these storms to strike the New Orleans area. 

The effect of storm heading angle on surges at the coast appears to be twofold. First, the 

overall along-coast pattern is broadened; since the storm moves along the coast at the same time 

that it moves toward landfall. Second, there is a relatively slow variation in the maximum surges 

produced by a storm as a function of the angle of the storm track with the coast; however, as 

seen in Figure D8, maximum surge is relatively weakly dependent on the angle of storm 

intersection with the coast. In general, the hurricane approaching slightly (15-30 degrees) from 

west of straight onto a straight east-west coast produce a somewhat higher surge (5% or so) than 

hurricanes moving perpendicularly to the coast. On the other hand, hurricanes approaching the 

straight east-west coast from a more easterly direction will tend to produce lower surges than 

produced by hurricanes moving perpendicular to such a coast. This appears to be a fairly broad 

pattern that can be represented via interpolation. 

The effect of forward storm speed is addressed by considering three different forward 

velocities Vf=(11,6,17) knots, where 11 is around the mean and the 6-kt and 17-kt speeds span 

almost the entire range of Vf values at landfall for storms with Cp’s less than 950. Note that a 

few of the track-angle combinations are dropped due to either their very oblique angle with 

respect to the shore or they were exiting storms from the New Orleans area moving toward the 
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northeast. Increased forward storm speed contributes to higher wind speeds in the hurricane PBL 

model. Consequently, one effect of increasing forward storm velocity is to increase the surge at 

the coast by a factor which similar to increasing the wind speeds within the hurricane, i.e. 

D5. 

1

2

2

max

1 2

max

1 1

2 2

max

0.5

0.5

where

 is the surge at the coast in storm 1, with forward speed = 

 is the surge at the coast in storm 2, with forward speed = 

 is the maximum wind speed 

f

f

v v

v v

v

v

v

η η

η

η

⎛ ⎞ + = ⎜ ⎟

+ ⎝ ⎠

th

of a stationary storm

 is the forward storm velocity of the i storm i f v

A second effect of storm speed is to change the duration that a flood wave has to propagate 

inland. Thus, a slowly moving storm may produce more extensive inland flooding than a faster 

moving storm. By covering essentially the entire range of forward storm speeds observed in 

major storms within the Gulf (see Figures 11a and 11b), we should be able to quantify the range 

of the effects of storm speed on surges in the New Orleans area. 

The information below provides information on the variation parameter combinations used in 

the New Orleans 152-storm study. 

Primary Tracks 

 Track 

(Vf=11) 1 2 3 4 4’ 5

Mean angle 9 9 9 9 9 (use Cp-Rmax combination set A) – 

 storms 1-45 

-45 4 4 4 4 (use Cp-Rmax combination set B) – 

 storms 46-61 

+45 4 4 4 4 (use Cp-Rmax combination set B) – 

 storms 66-81 

TOTAL = WAS 81 — NOW 77 

 Track 

(Vf= 6) 1 2 3 4 4’ 5

Mean angle 2 2 2 2 2 (use Cp-Rmax combination set C) – 

 storms 82-91 

-45 1 1 1 1 (use Cp-Rmax combination set D) – 

 storms 92-95 

+45 1 1 1 1 (use Cp-Rmax combination set D) – 

 storms 97-100 

TOTAL = 19 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-88

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 Track 

(Vf=17) 1 2 3 4 4’ 5

Mean angle 1 1 1 1 1 (use Cp-Rmax combination set D) – 

 storms 101-105 

-45 1 1 1 1 (use Cp-Rmax combination set D) – 

 storms 106-109 

+45 1 1 1 1 (use Cp-Rmax combination set D) – 

 storms 111-114 

TOTAL = WAS 14 – NOW 13 

Secondary Tracks 

 Track 

(Vf = 11) 1b 2b 3b 4b

Mean angle 2 2 2 2 (use Cp-Rmax combination set C) – storms 115-122 

-45 2 2 2 (use Cp-Rmax combination set C) – storms 123-128 

+45 2 2 2 (use Cp-Rmax combination set C) – storms 131-136 

TOTAL = 20 

 Track 

(Vf = 6) 1b 2b 3b 4b

Mean angle 2 2 2 2 (use Cp-Rmax combination set C) – storms 137-144 

-45 1 1 1 (use Cp-Rmax combination set D) – storms 145-147 

+45 1 1 1 (use Cp-Rmax combination set D) – storms 149-151 

TOTAL = 14 

 Track 

(Vf = 17) 1b 2b 3b 4b

Mean angle 1 1 1 1 (use Cp-Rmax combination set D) – storms 152-155 

-45 1 1 1 (use Cp-Rmax combination set D) – storms 156-158 

+45 1 1 1 (use Cp-Rmax combination set D) – storms 160-162 

TOTAL = 10 

GRAND TOTAL = 152 
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Figure D1. Distribution of normalized maximum surges along the coast (local surge maximm (ζ) divided by 

the maximum surge within the entire storm (ζmax) versus normalized distance along the coast 

(distance from storm peak divided by radius scaling parameter, Rmax, for offshore slopes in the 

range of 1:250 to 1:5000. Results are from numerical simulations on an idealized, straight 

coast for hurricanes for relatively small storm: Rmax (= Rp scaling radius in Cardone PBL model) 

= 10 nm. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-90

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure D2. Distribution of normalized maximum surges along the coast (local surge maximm (ζ) divided by 

the maximum surge within the entire storm (ζmax) versus normalized distance along the coast 

(distance from storm peak divided by radius scaling parameter, Rmax, for offshore slopes in the 

range of 1:250 to 1:5000. Results are from numerical simulations on an idealized, straight 

coast for hurricanes for relatively small storm: Rmax (= Rp scaling radius in Cardone PBL model) 

= 20 nm. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-91

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure D3. Distribution of normalized maximum surges along the coast (local surge maximm (ζ) divided by 

the maximum surge within the entire storm (ζmax) versus normalized distance along the coast 

(distance from storm peak divided by radius scaling parameter, Rmax, for offshore slopes in the 

range of 1:250 to 1:5000. Results are from numerical simulations on an idealized, straight 

coast for hurricanes for relatively small storm: Rmax (= Rp scaling radius in Cardone PBL model) 

= 30 nm. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-92

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure D4. Distribution of normalized storm surge (local maximum ([η(x)] divided by maximum over entire 

storm (ηmax) as a function of normalized distance along the coast. 
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A. 1:10000 B. 

2.5

3.0

3.5

4.0

4.5

5.0

5.5

6.0

6.5

7.0

7.5

2.5 3.0 3.5 4.0 4.5 5.0 5.5 6.0 6.5 7.0 7.5

Peak Surge for Straight Shoreline (m)

Peak Surge for Idealized New Orleans Shoreline (m)

C. 

Figure D5(A-C). Figure D5-A shows an idealized 

representation of the New Orleans 

coastal area, with a section of land 

protruding from a generalized straightline coast. Figure D5-B shows that the 

resulting surge distributions along the 

coast. Figure D5-C shows that surge 

values for this coastal configuration tend 

to be about 10-20% higher than the 

corresponding surges on a straight-line 

coast. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-94

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure D6. Figures D1-D3 re-plotted with lines 

approximately 31 nm drawn for the case 

in which the peak falls midway between 

the two tracks. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-95

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure D7. Sample time plot of the variation in central pressure, maximum wind speed, Rp, and forward 

storm speed used in quality control check of storm parameter behavior for first storm in JPM 

sequence. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-96

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure D8. Variation in maximum storm surge produced by hurricanes approaching a straight, shallowsloping (1:10,000) coast relative to the maximum surge produced by a storm moving 

perpendicularly to the coast. 

So = 1:10000, Vf = 10 kt, Track Variation

0.00

1.00

2.00

3.00

4.00

5.00

6.00

7.00

0.00 1.00 2.00 3.00 4.00 5.00 6.00 7.00

Surge, Track Due North (m)

Surge, Track at Angle Specified in Legend (m)

15

30

45

-15

-30

-45

match

-5% match

+5% match
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-97

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

APPENDIX E of R2007 

The Characteristics of the Holland Pressure Profile Parameter 

and the Radius to Maximum Winds for Gulf of Mexico and 

Atlantic Hurricanes as Determined from an Analysis of Flight 

Level Pressure Data and H*Wind Surface Wind Speed Data 

Introduction 

Using pressure data collection during hurricane reconnaissance flights, coupled with 

additional information derived from the Hurricane Research Division’s H*Wind snapshots of 

hurricane wind fields, an analysis of the radius to maximum winds and the Holland B parameter 

was performed. The reconnaissance data incorporates flights encompassing the time period 1977 

through to 2001, but the analysis was limited to include only those data collected at the 700 mbar 

or higher level. 

The Holland B parameter was found to be inversely correlated with both the size of a 

hurricane and the latitude of a hurricane. A weak positive correlation of B with central pressure 

deficit and sea surface temperature was also observed. A statistical relationship between the 

Holland B parameter and a non-dimensional parameter incorporating central pressure, radius to 

maximum winds, sea surface temperature and latitude was developed. 

A qualitative examination of the variation of B, central pressure and radius to maximum 

winds as a function of time suggests that along the Gulf of Mexico coastline (excluding 

Southwest Florida), during the final 6 hrs to 24 hrs before landfall the hurricanes weaken as 

characterized by both an increase in central pressure and the radius to maximum winds, and a 

decrease in the Holland B parameter. This weakening characteristic of landfalling storms was not 

as evident for hurricanes making landfall elsewhere along the United States coastline. 

Flight Level Data Analysis Methodology 

Upper level aircraft data available at NOAA site were used to estimate Holland’s pressure 

profile parameter (B). The upper level aircraft dataset used here contains a total of 4546 radial 

profiles from 62 Atlantic storms. For every storm, data has been organized based on the different 

flights that passed through the storm. For each flight, the airplane traversed through the hurricane 

a number of times in different directions. For every pass the data was collected from the center 

of the storm to a certain radius (usually 150 km). Available data is then organized according to 

their radial distance from the center of the storm. For each bin (based on the radius from the 

center of the storm) flight level pressure, flight altitude, dew point temperature, wind speed and 

air temperature are available. Each profile from every flight and every storm is treated as an 

independent observation. Holland, (1980) describes the radial distribution of surface pressure in 

a hurricane in the form: 

⎥

⎦

⎤ ⎢

⎣

⎡ = + Δ ⋅ − B

r

A

p(r) p0 p exp (1) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-98

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where p(r) is the surface pressure at a distance r from the storm center, p0 is the central pressure, 

Δp is the central pressure difference, A is the location parameter and B is the Holland’s pressure 

profile parameter. Holland (1980) showed that RMW = A1/B where RMW is the radius to 

maximum winds, and thus Equation (1) can be expressed as: 

B

r

RMW

p r p p ⎥

⎦

⎤ ⎢

⎣

⎡ ( ) = 0 + Δ ⋅ exp− (2) 

The surface pressure and radial distance are transformed to the form of Equation (2). The 

missing quantities in Equation (2) are RMW and B. First estimate of RMW is made from the 

recorded wind speed profile i.e. RMW is the radius to the measured maximum wind speed. From 

here on, the radius corresponding to the maximum wind speed in a profile is referred to as RMW. 

To estimate the optimum values of B and RMW, RMW and B are varied over the range [0.5RMW, 

1.5RMW] and [0.5, 2.5] respectively. The algorithm calculates an optimum B value by 

minimizing the mean of the square differences between the measured and the modeled surface 

pressure in a range of 0.5RMW to 1.5RMW for different B and RMW values. Mathematically, the 

mean square error between the measured and the modeled surface pressure can be written as: 

( ) 1.5 2

2 0.5 i i

RMW

obs theo

i RMW

P P

n

ε =

−

=

∑

 (3) 

where Pobsi is the measured pressure, Ptheoi is the theoretical pressure calculated using equation 

(2) and n is the number of data points in the range [0.5RMW, 1.5RMW]. The values of B and

RMW chosen correspond to those yielding the minimum mean square error, ε

2

. The 

corresponding r

2

 value for the fit is given by: 

2

2

2 1

σ

ε

r = − (4) 

where σ is the standard deviation of the measured pressure data in the range of [0.5RMW, 

1.5RMW]. 

Quality control criteria 

A quality control criterion was used to filter out profiles. Each of the filtered profiles has at 

least one of the following characteristics associated with it, (a) Flight level pressure is less than 

700 mbar i.e. height greater than 3000 m, (b) Central pressure difference is less than 25 mbar, 

(c) Radius to maximum winds is greater than two-third of the sampling domain, (d) the distance 

of aircrafts closest approach to the center is greater than half of the radius to maximum winds, 

(e) Data is available for less than one third of the sampling range i.e. less than 50 km, (f) visual 

inspection which involved eliminating profiles with a considerable amount of data missing in the 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-99

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range of interest [0.5RMW, 1.5RMW ]. The rationale for using criteria (a) is that higher the 

measurement height, less representative measurements are of the surface observations. Criteria 

(b) results in the data associated with Category 1 or higher hurricanes only. The rationale for 

using criteria (c), (d), (e) and (f) is to ensure that there is a sufficient number of measurements on 

both sides of the radius to maximum winds to have a clear representation of the shape of the 

profile. 

The use of the quality control criteria eliminated a total of 2291 profiles from a set of 

4556 profiles. Table 1 presents the count of the eliminated pressure profiles based on the filtering 

criteria. It is clear that criteria (a) and (b) are the most common reasons for profile elimination. 

Storm by storm percentage of the retained profiles is given in Table 2. For some storms, no 

profiles were retained as all the profiles either had a central pressure difference of less than 

25 mbar (e.g. Chantal 1995) or a flight level pressure of less than 700 mbar (e.g. Hugo 1989). 

Figure 1 presents a few examples of pressure profiles that were eliminated from the analysis. 

Both the measured pressure data and the corresponding fit to Holland’s equation are shown. It is 

observed that each of the subplots in Figure 1 is compromised by at least one of the above 

mentioned quality control criteria. 

Figure 2 presents examples of pressure profiles that were retained for analysis. Each row in 

Figure 2 corresponds to a complete airplane traverse in one direction. The shaded regions in 

Figure 2 represent the error minimizing range of 0.5RMW to 1.5RMW. The fit parameters i.e. the 

B value, the central pressure difference and the RMW are also provided in the title of every 

profile. For a given traverse through a hurricane, differences in the B values for two different 

profiles is due to the change in the radius to the maximum winds and the central pressure 

difference. The geographical distribution of the filtered profiles, based on the storm center, is 

shown in Figure 3. The filtered profiles have a wide geographical distribution and provides with 

a wide domain of hurricane climatic characteristics. The filtered dataset has an average RMW of 

46 km (standard deviation of 22 km), an average central pressure difference of 51 mbar 

(18 mbar) and an average location of 25.84°N (5.74°N) and 74.78°W (12.82°W). 71% of the fits 

yield r

2

 values greater than 0.95 and 80% of the fits have a mean square error less than 2.5 mbar. 

The maximum mean square error was 24.6 mbar which occurred for one of Hurricane Opal’s 

profiles where Holland’s equation overestimated the pressures at all points. 

Table 1. 

Distribution of filtered pressure profiles based on filtering criteria. 

Filter criteria Number of profiles eliminated 

(a) 459 

(b) 1180 

(c) 121 

(d)+(e)+(f) 531 

Total number of filtered profiles 2291 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-100

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Table 2. 

Percentage of flight level pressure profiles retained. 

Storm Year Total Retained %Retained Comments 

no-name 1938 5 5 100.00 Data extracted manually from Myers & Jordan (1956) 

Anita 1977 20 20 100.00 

David 1979 24 17 70.83 

Frederic 1979 62 38 61.29 

Allen 1980 125 43 34.40 

Gert 1981 78 1 1.28 Δp<25mb for all the cases, except one. 

Alicia 1983 50 39 78.00 

Arthur 1984 22 0 0.00 Δp <25mb for all the cases. 

Diana 1984 128 67 52.34 

Danny 1985 26 0 0.00 Δp <25mb for all the cases. 

Elena 1985 122 99 81.15 

Gloria 1985 42 24 57.14 

Isabel 1985 48 0 0.00 Δp <25mb for all the cases. 

Juan 1985 36 6 16.67 

Charley 1986 28 0 0.00 Δp <25mb for all the cases. 

Emily 1987 56 1 1.79 40 out of 56 profiles have flight level pressure <700mb. 

Floyd 1987 22 0 0.00 Δp <25mb for all the cases. 

Florence 1988 20 11 55.00 

Gilbert 1988 50 39 78.00 

Joan 1988 6 5 83.33 

Dean 1989 12 1 8.33 

Gabrielle 1989 12 10 83.33 

Hugo 1989 40 0 0.00 Flight level pressure <700mb for all the cases 

Jerry 1989 17 5 29.41 

Gustav 1990 84 82 97.62 

Bob 1991 92 34 36.96 

Claudette 1991 73 71 97.26 

Andrew 1992 141 95 67.38 

Debby 1994 10 0 0.00 Δp <25mb for all the cases. 

Gordon 1994 83 8 9.64 57 out of 83 profiles have Δp <25mb. 

Allison 1995 39 3 7.69 35 out of 39 profiles have Δp <25mb. 

Chantal 1995 72 0 0.00 Δp <25mb for all the cases. 

Erin 1995 97 66 68.04 

Felix 1995 130 59 45.38 

Gabrielle 1995 16 0 0.00 Δp <25mb for all the cases. 

Iris 1995 132 41 31.06 

Luis 1995 130 77 59.23 

Marilyn 1995 116 96 82.76 

Opal 1995 76 21 27.63 

Roxanne 1995 141 52 36.88 

Bertha 1996 78 56 71.79 

Cesar 1996 34 0 0.00 Δp <25mb for all the cases. 

Edouard 1996 178 135 75.84 

Fran 1996 143 102 71.33 

Hortense 1996 109 59 54.13 

Josephine 1996 23 1 4.35 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-101

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Storm Year Total Retained %Retained Comments 

Kyle 1996 8 0 0.00 Δp <25mb for all the cases. 

Lili 1996 68 28 41.18 

Marco 1996 67 1 1.49 Δp <25mb for all the cases, except two. 

Erika 1997 56 36 64.29 

Bonnie 1998 193 113 58.55 

Danielle 1998 133 48 36.09 

Earl 1998 32 3 9.38 

Georges 1998 202 125 61.88 

Mitch 1998 86 57 66.28 

Bret 1999 102 49 48.04 

Dennis 1999 158 83 52.53 

Floyd 1999 163 103 63.19 

Keith 2000 50 40 80.00 

Leslie 2000 29 0 0.00 Δp <25mb for all the cases. 

Michael 2000 21 11 52.38 

Humberto 2001 46 13 28.26 

Michelle 2001 89 61 68.54 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-102

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Figure 1. Examples of the eliminated profiles. 

982

988

995

Bertha 12/07/1996-08:30 

Pressure (mb)

983

989

995

Josephine 07/10/1996-11:35

981

987

994

Michelle 03/11/2001-17:24 

Pressure (mb)

953

978

1004

Floyd 13/10/1987-02:15 

958

983

1007

Floyd 16/09/1999-03:42 

Pressure (mb)

964

988

1012

Claudette 07/09/1991-17:28

987

998

1008

Bertha 12/07/1996-01:12 

Pressure (mb)

990

997

1003

Earl 03/09/1998-00:45 

0 50 100 150

943

972

1002

Allen 05/08/1980-16:15 

Radius (km)

Pressure (mb)

0 50 100 150

952

980

1008

Andrew 26/08/1992-02:10 

Radius (km)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-103

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Figure 2. Examples of surface pressure profiles for a traverse across a given hurricane. 

150 100 50 0

940

960

980

1000

1020

Andrew -1992/08/23-0542, B=1.6,Δp=75mb,Rmax=11km 

Radius(km)

Pressure(mb)

0 50 100 150

940

960

980

1000

1020

Andrew -1992/08/23-0550, B=1.7,Δp=79mb,Rmax=13km 

Radius(km)

Pressure(mb)

150 100 50 0

970

980

990

1000

1010

1020

1030

Bertha-1996/07/09-0647, B=1.45, Δp=52mb, Rmax=35km 

Radius(km)

Pressure(mb)

0 50 100 150

970

980

990

1000

1010

1020

1030

Bertha-1996/07/09-0658, B=1.25, Δp=53mb, Rmax=32km 

Radius(km)

Pressure(mb)

150 100 50 0

970

980

990

1000

1010

1020

1030

Bret-1999/08/21-2110, B=1.7, Δp=52mb, Rmax=22km 

Radius(km)

Pressure(mb)

0 50 100 150

970

980

990

1000

1010

1020

1030

Bret-1999/08/21-2118, B=1.9, Δp=53mb, Rmax=20km 

Radius(km)

Pressure(mb)

150 100 50 0

940

960

980

1000

1020

Edouard-1996/08/26-1855, B=1.35, Δp=74mb, Rmax=22.5km

Radius(km)

Pressure(mb)

0 50 100 150

960

980

1000

1020

Edouard-1996/08/26-1904, B=1.65, Δp=70mb, Rmax=30km 

Radius(km)

Pressure(mb)
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Figure 2. (continued) Examples of surface pressure profiles for a traverse across a given hurricane. 

150 100 50 0

940

960

980

1000

1020

Floyd-1999/09/15-0518, B=1.15, Δp=80mb, Rmax=66.5km 

Radius(km)

Pressure(mb)

0 50 100 150

940

960

980

1000

1020

Floyd-1999/09/15-0525, B=1.1, Δp=81mb, Rmax=65km 

Radius(km)

Pressure(mb)

150 100 50 0

980

990

1000

1010

1020

1030

Georges-1998/09/26-1108, B=1, Δp=42mb, Rmax=53.5km 

Radius(km)

Pressure(mb)

0 50 100 150

980

990

1000

1010

1020

1030

Georges-1998/09/26-1116, B=1.4, Δp=42mb, Rmax=57.5km 

Radius(km)

Pressure(mb)

150 100 50 0

970

980

990

1000

1010

1020

1030

Gustav-1990/08/27-1856, B=1.55, Δp=47mb, Rmax=23.5km 

Radius(km)

Pressure(mb)

0 50 100 150

970

980

990

1000

1010

1020

1030

Gustav-1990/08/27-1859, B=1.15, Δp=46mb, Rmax=22.5km 

Radius(km)

Pressure(mb)

150 100 50 0

960

980

1000

1020

Luis-1995/09/04-1700, B=1.4, Δp=71mb, Rmax=42.5km 

Radius(km)

Pressure(mb)

0 50 100 150

960

980

1000

1020

Luis-1995/09/04-1704, B=1.5, Δp=71mb, Rmax=38.5km 

Radius(km)

Pressure(mb)
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Figure 3. Geographical distribution of all the filtered profiles. 

The approach for analyzing the B and RMW data involved the estimation of RMW and B

from each single pass of a flight through the storm, and then smoothing the variations in B and 

RMW as a function of time. Figure 4 presents ten examples of both the single flight (point 

estimates) and the smoothed estimates of B and RMW plotted vs. time, for landfalling hurricanes. 

The landfall time is indicated with a vertical line in each plot. Using the smoothed data, values of 

B and RMW were extracted at intervals of approximately 3 hrs and retained for use in the 

statistical analyses. The mean values of B and RMW for the smoothed data set are 1.21 and 

47 km respectively. The corresponding standard deviations are 0.29 and 21 km respectively. 

Note that in only one of the 11 landfall’s indicated in Figure 4, does the Holland B parameter 

appear to increase as a hurricane approaches land (Hurricane Floyd near the NC coast). Table 3 

summarizes, qualitatively, the tendency in the changes of B over the final few hours before 

landfall. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-106

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Figure 4. Examples of Smoothed (line) and Point Estimates (symbols) of RMW and B derived from 

700 mbar level pressure data. Vertical line(s) represent time of landfall. 

Andrew, 1992

0.0

0.5

1.0

1.5

2.0

2.5

8/22/92

0:00

8/23/92

0:00

8/24/92

0:00

8/25/92

0:00

8/26/92

0:00

8/27/92

0:00

Date (UTC)

Holland B Parameter

Andrew, 1992

0

5

10

15

20

25

30

35

8/22/92

0:00

8/23/92

0:00

8/24/92

0:00

8/25/92

0:00

8/26/92

0:00

8/27/92

0:00

Date (UTC)

RMW (km)

Elena, 1985

0.0

0.5

1.0

1.5

2.0

2.5

8/29/85

0:00

8/30/85

0:00

8/31/85

0:00

9/1/85

0:00

9/2/85

0:00

9/3/85

0:00

Date (UTC)

Holland B Parameter

Elena, 1985

0

20

40

60

80

100

8/29/85

0:00

8/30/85

0:00

8/31/85

0:00

9/1/85

0:00

9/2/85

0:00

9/3/85

0:00

Date (UTC)

RMW (km)

Opal, 1995

0.0

0.5

1.0

1.5

2.0

2.5

10/2/95

0:00

10/2/95

12:00

10/3/95

0:00

10/3/95

12:00

10/4/95

0:00

10/4/95

12:00

10/5/95

0:00

10/5/95

12:00

Date (UTC)

Holland B Parameter

Opal, 1995

0

20

40

60

80

100

120

10/2/95

0:00

10/2/95

12:00

10/3/95

0:00

10/3/95

12:00

10/4/95

0:00

10/4/95

12:00

10/5/95

0:00

10/5/95

12:00

Date (UTC)

RMW (km)

Frederic, 1979

0.0

0.5

1.0

1.5

2.0

2.5

9/11/79

0:00

9/11/79

12:00

9/12/79

0:00

9/12/79

12:00

9/13/79

0:00

9/13/79

12:00

Date (UTC)

Holland B Parameter

Frederic, 1979

0

10

20

30

40

50

60

70

9/11/79

0:00

9/11/79

12:00

9/12/79

0:00

9/12/79

12:00

9/13/79

0:00

9/13/79

12:00

Date (UTC)

RMW (km)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-107

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Figure 4. (continued) Examples of Smoothed (line) and Point Estimates (symbols) of RMW and B

derived from 700 mbar level pressure data. Vertical line(s) represent time of landfall. 

Georges, 1998

0.0

0.5

1.0

1.5

2.0

2.5

9/18/98

0:00

9/20/98

0:00

9/22/98

0:00

9/24/98

0:00

9/26/98

0:00

9/28/98

0:00

9/30/98

0:00

Time (UTC)

Holland B Parameter

Georges, 1998

0

20

40

60

80

100

9/18/98

0:00

9/20/98

0:00

9/22/98

0:00

9/24/98

0:00

9/26/98

0:00

9/28/98

0:00

9/30/98

0:00

Date (UTC)

RMW (km)

Fran, 1996

0

0.5

1

1.5

2

2.5

8/31/96

0:00

9/1/96

0:00

9/2/96

0:00

9/3/96

0:00

9/4/96

0:00

9/5/96

0:00

9/6/96

0:00

9/7/96

0:00

Date (UTC)

Holland B Parameter

Fran, 1996

0

20

40

60

80

100

8/31/96

0:00

9/1/96

0:00

9/2/96

0:00

9/3/96

0:00

9/4/96

0:00

9/5/96

0:00

9/6/96

0:00

9/7/96

0:00

Date (UTC)

RMW (km)

Bertha, 1996

0.0

0.5

1.0

1.5

2.0

2.5

7/8/96

0:00

7/9/96

0:00

7/10/96

0:00

7/11/96

0:00

7/12/96

0:00

7/13/96

0:00

Date (UTC)

Holland B Parameter

Bertha, 1996

0

20

40

60

80

100

7/8/96 0:00 7/9/96 0:00 7/10/96

0:00

7/11/96

0:00

7/12/96

0:00

7/13/96

0:00

Date (UTC)

RMW (km)

Bonnie, 1998

0.0

0.5

1.0

1.5

2.0

2.5

8/22/98

0:00

8/23/98

0:00

8/24/98

0:00

8/25/98

0:00

8/26/98

0:00

8/27/98

0:00

8/28/98

0:00

8/29/98

0:00

Date (UTC)

Holland B Parameter

Bonnie, 1998

0

20

40

60

80

100

8/22/98

0:00

8/23/98

0:00

8/24/98

0:00

8/25/98

0:00

8/26/98

0:00

8/27/98

0:00

8/28/98

0:00

8/29/98

0:00

Date (UTC)

RMW (km)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-108

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 4. (concluded) Examples of Smoothed (line) and Point Estimates (symbols) of RMW and B

derived from 700 mbar level pressure data. Vertical line(s) represent time of landfall. 

Table 3. 

Tendency of Holland B Parameter for Landfalling Storms 

Hurricane and Landfall Location B Tendency at landfall

Frederic (Alabama) ~ constant 

Elena (Mississippi) ~ constant 

Andrew South Florida ~constant to ~negative 

Andrew Louisiana negative 

Opal (North West Florida) constant 

Bertha (North Carolina) negative 

Fran (North Carolina) ~constant 

Bonnie (North Carolina) negative 

Georges (Mississippi) negative 

Bret (Texas) ~constant 

Floyd (North Carolina) positive 

Floyd, 1999

0.0

0.5

1.0

1.5

2.0

2.5

9/10/99

0:00

9/11/99

0:00

9/12/99

0:00

9/13/99

0:00

9/14/99

0:00

9/15/99

0:00

9/16/99

0:00

9/17/99

0:00

Date (UTC)

Holland B Parameter

Floyd, 1999

0

20

40

60

80

100

120

9/10/99

0:00

9/11/99

0:00

9/12/99

0:00

9/13/99

0:00

9/14/99

0:00

9/15/99

0:00

9/16/99

0:00

9/17/99

0:00

Date (UTC)

RMW (km)

Bret, 1999

0.0

0.5

1.0

1.5

2.0

2.5

8/20/99

12:00

8/21/99

0:00

8/21/99

12:00

8/22/99

0:00

8/22/99

12:00

8/23/99

0:00

8/23/99

12:00

Date (UTC)

Holland B Parameter

Bret, 1999

0

5

10

15

20

25

30

35

8/20/99

12:00

8/21/99

0:00

8/21/99

12:00

8/22/99

0:00

8/22/99

12:00

8/23/99

0:00

8/23/99

12:00

Date (UTC)

RMW (km)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-109

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Supplemental H*Wind Data 

The flight level data encompasses storms through to 2001, and thus to supplement the data 

set with more recent storms, some additional storms analyzed using the H*Wind methodology 

were added. The only storms added were the intense storms from the 2004 and 2005 seasons that 

had been re-analyzed using the most recent SFMR calibrations. The intense storms that have 

been reanalyzed include Hurricanes Katrina (2005) and Hurricane Ivan (2004). Hurricane Rita 

was added to the data set even though it had not been re-analyzed, because at its most intense, 

the storm had a minimum central pressure of less than 900 mbar. Using the wind field model 

described in Vickery, et al. (2007)., and the values of central pressure, RMW, storm translation 

speed, and the maximum sustained wind speed, a B value chosen so that the maximum surface 

level wind speed (one minute sustained value) obtained from the model match the H*Wind 

estimate of the maximum wind speed. Thus the estimated B values are obtained through an 

indirect measure, matching the maximum wind speed rather than the shape of the entire wind 

field. 

Figure 5 presents plots of RMW, maximum one minute surface level wind speed, the derived 

B parameter and central pressure as a function of time for the three aforementioned hurricanes, in 

addition to the data derived for hurricanes Dennis (2005), Bret (1999) and Lili (1999). These 

three additional storms are given to examine the change in the characteristics of the storms as 

they approach land. Each plot also presents the central pressure(s) at land fall as given in the 

NHC hurricane reports. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-110

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 5. Smoothed (line) and Point Estimates (symbols) of RMW and B derived from H*Wind data. 

Vertical line(s) represent time of landfall. Solid square point at time of landfall represents NHC 

landfall pressure value. 

Hurricane Katrina (2005)

0.0

0.5

1.0

1.5

2.0

2.5

8/26/05 0:00 8/27/05 0:00 8/28/05 0:00 8/29/05 0:00 8/30/05 0:00

Time (UTC)

Holland B Parameter

Hurricane Katrina (2005)

0

10

20

30

40

50

60

70

8/26/05 0:00 8/27/05 0:00 8/28/05 0:00 8/29/05 0:00 8/30/05 0:00

Time (UTC)

RMW (km)

Hurricane Katrina (2005)

0

20

40

60

80

100

120

140

160

180

8/26/05 0:00 8/27/05 0:00 8/28/05 0:00 8/29/05 0:00 8/30/05 0:00

Time (UTC)

Maximum One Minute Wind 

Speed (mph)

Hurricane Katrina (2005)

900

910

920

930

940

950

960

970

980

990

1000

8/26/05 0:00 8/27/05 0:00 8/28/05 0:00 8/29/05 0:00 8/30/05 0:00

Time (UTC)

Central Pressure (mbar)

Hurricane Rita (2005)

880

900

920

940

960

980

1000

9/20/05

0:00

9/21/05

0:00

9/22/05

0:00

9/23/05

0:00

9/24/05

0:00

9/25/05

0:00

Time (UTC)

Central Pressure (mbar)

Hurricane Rita(2005)

0

20

40

60

80

100

120

140

160

180

9/20/05

0:00

9/21/05

0:00

9/22/05

0:00

9/23/05

0:00

9/24/05

0:00

9/25/05

0:00

Time (UTC)

Maximum One Minute Wind 

Spped (mph)

Hurricane Rita (2005)

0

5

10

15

20

25

30

35

40

9/20/05

0:00

9/21/05

0:00

9/22/05

0:00

9/23/05

0:00

9/24/05

0:00

9/25/05

0:00

Time (UTC)

RMW (km)

Hurricane Rita (2005)

0.0

0.5

1.0

1.5

2.0

2.5

9/20/05

0:00

9/21/05

0:00

9/22/05

0:00

9/23/05

0:00

9/24/05

0:00

9/25/05

0:00

Time (UTC)

Holland B Parameter
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-111

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 5. (continued) Smoothed (line) and Point Estimates (symbols) of RMW and B derived from 

H*Wind data. Vertical line(s) represent time of landfall. Solid square point at time of landfall 

represents NHC landfall pressure value. 

Hurricane Dennis (2005)

0

5

10

15

20

25

30

35

40

7/6/05 0:00 7/7/05 0:00 7/8/05 0:00 7/9/05 0:00 7/10/05

0:00

7/11/05

0:00

Time (UTC)

RMW (km)

Hurricane Dennis (2005)

0.0

0.5

1.0

1.5

2.0

2.5

7/6/05 0:00 7/7/05 0:00 7/8/05 0:00 7/9/05 0:00 7/10/05

0:00

7/11/05

0:00

Time (UTC)

Holland B Parameter

Hurricane Dennis (2005)

0

20

40

60

80

100

120

140

160

7/6/05

0:00

7/7/05

0:00

7/8/05

0:00

7/9/05

0:00

7/10/05

0:00

7/11/05

0:00

Time (UTC)

Maximum One Minute Wind 

Speed (mph)

Hurricane Dennis (2005)

920

930

940

950

960

970

980

990

1000

7/6/05

0:00

7/7/05

0:00

7/8/05

0:00

7/9/05

0:00

7/10/05

0:00

7/11/05

0:00

Time (UTC)

Central Pressure (mbar)

Hurricane Ivan (2004)

0

20

40

60

80

100

120

140

160

180

9/7/04

0:00

9/9/04

0:00

9/11/04

0:00

9/13/04

0:00

9/15/04

0:00

9/17/04

0:00

Time (UTC)

Maximum One Minute Wind 

Speed (mph)

Hurricane Ivan (2004)

0

5

10

15

20

25

30

35

40

45

9/7/04

0:00

9/9/04

0:00

9/11/04

0:00

9/13/04

0:00

9/15/04

0:00

9/17/04

0:00

TIme (UTC)

RMW (km)

Hurricane Ivan (2004)

0.0

0.5

1.0

1.5

2.0

2.5

9/7/04

0:00

9/9/04

0:00

9/11/04

0:00

9/13/04

0:00

9/15/04

0:00

9/17/04

0:00

Time (UTC)

Holland B Parameter

Hurricane Ivan (2004)

900

910

920

930

940

950

960

970

980

9/7/04

0:00

9/9/04

0:00

9/11/04

0:00

9/13/04

0:00

9/15/04

0:00

9/17/04

0:00

Time (UTC)

Central pressure (mbar)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-112

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 5. (concluded) Smoothed (line) and Point Estimates (symbols) of RMW and B derived from 

H*Wind data. Vertical line(s) represent time of landfall. Solid square point at time of landfall 

represents NHC landfall pressure value. 

Hurricane Lili, 2002

0

20

40

60

80

100

120

140

160

9/29/02

0:00

9/30/02

0:00

10/1/02

0:00

10/2/02

0:00

10/3/02

0:00

10/4/02

0:00

Time (UTC)

Maximum One Minute Wind 

Speed (mph)

Hurricane Lili, 2002

0.0

0.5

1.0

1.5

2.0

2.5

9/29/02

0:00

9/30/02

0:00

10/1/02

0:00

10/2/02

0:00

10/3/02

0:00

10/4/02

0:00

Time (UTC)

Holland B Parameter

Hurricane Lili, 2002

0

10

20

30

40

50

9/29/02

0:00

9/30/02

0:00

10/1/02

0:00

10/2/02

0:00

10/3/02

0:00

10/4/02

0:00

Time (UTC)

RMW (km)

Hurricane Lili, 2002

930

940

950

960

970

980

990

1000

1010

9/29/02

0:00

9/30/02

0:00

10/1/02

0:00

10/2/02

0:00

10/3/02

0:00

10/4/02

0:00

Time (UTC)

Central Pressure (mbar)

Hurricane Bret (1999)

0

5

10

15

20

25

30

8/22/99 6:00 8/22/99 12:00 8/22/99 18:00 8/23/99 0:00 8/23/99 6:00

Time (UTC)

RMW (km)

Hurricane Bret (1999)

0.0

0.5

1.0

1.5

2.0

2.5

8/22/99 6:00 8/22/99 12:00 8/22/99 18:00 8/23/99 0:00 8/23/99 6:00

Time (UTC)

Holland B Parameter

Hurricane Bret (1999)

940

950

960

970

980

8/22/1999

6:00

8/22/1999

12:00

8/22/1999

18:00

8/23/1999

0:00

8/23/1999

6:00

Time (UTC)

Central Pressure (mbar)

Hurricane Bret (1999)

0

20

40

60

80

100

120

140

8/22/1999

6:00

8/22/1999

12:00

8/22/1999

18:00

8/23/1999

0:00

8/23/1999

6:00

Time (UTC)

Maximum One Minute Wind 

Speed (mph)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-113

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

These six hurricanes represent all the Gulf of Mexico landfalling hurricanes in the H*Wind 

database that include information on both wind speeds and central pressure in each of the 

H*Wind snapshots. Additional storms are given in the H*Wind database that do not have central 

pressures provided on the H*Wind snapshots. All of the six hurricanes show an increase in 

central pressure and a decrease in the magnitude of the Holland B parameter as they approach the 

Gulf Coast. An increase in the radius to maximum winds as the hurricanes approach landfall is 

also evident in five of the six cases examined. 

A similar analysis of hurricane characteristics for hurricanes making landfall in regions other 

than along the Gulf of Mexico coast did not indicate that there is a strong tendency for the storms 

to weaken and enlarge before landfall. 

Statistical Model for Radius to Maximum Winds 

All Hurricanes. The RMW for all points (flight level data plus H*Wind data) in the data set 

having a central pressure of less than 980 mbar were modeled as a function of central pressure 

difference and latitude in the form: 

5 2 4 2 ln( ) = 3.559 − 7.229×10 Δ + 5.296×10 Ψ − − RMW p ; r

2

=0.266, σlnRMW = 0.449 (4) 

An analysis of the errors (difference between the regression model estimates and the data) 

indicates that the model error reduced with increasing Δp, as indicated in Figure 6. 

Figure 6. Absolute value of RMW model error vs. Δp

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

30 40 50 60 70 80 90 100 110 120 130

Central Pressure Difference (mbar)

Absolute Value of Model Error

Absolute Value of Error

Absolute Value of Error Running Mean

Std Deviation

Modeled Standard Deviation
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-114

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The error, σlnRMW, is modeled in the form: 

σlnRMW = 0.460 Δp ≤ 87 mbar (5a) 

σlnRMW = 1.1703 – 0.00817Δp 87 mbar ≤ Δp ≤ 120 mbar (5b) 

σlnRMW = 0.190 Δp >120 mbar (5c) 

Figure 7 presents the modeled and observed values of RMW plotted vs. Δp. The modeled 

data are given as the median estimates and the range defined by ±2σlnRMW. The modeled range 

reflects the reduction in σlnRMW as a function of Δp. 

Figure 7. Modeled and observed RMW vs. Δp for all hurricanes 

Gulf of Mexico Hurricanes. In order to determine if the characteristics of the RMW 

associated with the Gulf of Mexico storms differed from that obtained using the all storm data, 

the RMW –Δp and RMW-ψ relationships were re-examined. For this analysis the Gulf of Mexico 

storms included all hurricanes west of 81o

W and north of 18oN. The RMW for all storms (flight 

level data plus H*Wind data) in the Gulf of Mexico data set with central pressures less than 980 

mbar were modeled as a function of central pressure difference in the form: 

5 2 ln(RMW ) = 3.859 − 7.700×10 Δp − r

2

=0.290, σlnRMW = 0.390 (6) 

0

25

50

75

100

125

150

20 40 60 80 100 120 140

Central Pressure Difference (mbar)

Radius to Maximum Winds (km)

Data

Model Median

+/- 2 Standard Deviations
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-115

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The RMW was found to be independent of latitude. As in the all storm case, the model error 

reduces with increasing Δp, as indicated in Figure 8. 

Figure 8. Absolute value of RMW model error vs. Δp for Gulf of Mexico hurricanes

The error, σlnRMW, for Gulf of Mexico hurricanes is modeled in the form: 

σlnRMW = 0.396 Δp ≤ 100 mbar (7a) 

σlnRMW = 1.424 – 0.01029Δp 100 mbar ≤ Δp ≤ 120 mbar (7b) 

σlnRMW = 0.19 Δp >120 mbar (7c) 

Figure 9 presents the modeled and observed values of RMW plotted vs. Δp for the Gulf of 

Mexico hurricanes. The modeled data are given as the median estimates and the range defined by 

±2σlnRMW. The modeled range reflects the reduction in σlnRMW as a function of Δp. 

Figure 10 presents the median values of the RMW computed using Equation 4 (all hurricane 

RMW model) computed for latitudes of 25o

N (Southern Gulf of Mexico) and 30oN (Northern 

Gulf of Mexico), where it is seen that for the Northern Gulf of Mexico storms, the all hurricanes 

RMW model over estimates the size of the Gulf of Mexico hurricanes, indicating that Gulf of 

Mexico hurricanes, are smaller than Atlantic hurricanes. 

0.0

0.2

0.4

0.6

0.8

1.0

1.2

30 40 50 60 70 80 90 100 110 120 130

Central Pressure Difference (mbar)

Absolute Value of Model Error

Absolute Value of Model error

Absolute Value of Model Error Running Average

Std. Deviation of Model Error

Modeled Standard Deviation
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-116

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 9. Modeled and observed RMW vs. Δp for Gulf of Mexico hurricanes. 

Figure 10. Comparison of all hurricanes model predicted median RMW to Gulf of Mexico model median 

RMW. 

0

20

40

60

80

100

120

20 40 60 80 100 120 140

Central Pressure Difference (mbar)

Radius to Maximum Winds (km)

Data

Model Median

Model +/- 2 Std Dev

0

10

20

30

40

50

60

20 40 60 80 100 120 140

Central Pressure Difference (mbar)

Radius to Maximum Winds (km)

Gulf of Mexico Model Median

All storms Model Median at 25 Degrees N

All storms Model Median at 30 Degrees N
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-117

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

RMW for Landfalling Storms 

Figure 11 presents the values of the RMW for storms making landfall along the Gulf and 

Atlantic coasts of the United States. In the case of Gulf Coast storms, no statistically significant 

correlation exists between the RMW and either latitude or Δp. In the case of hurricanes making 

landfall along the Atlantic coast, the RMW is positively correlated with latitude, and negatively 

correlated with the Δp

2

. As a group (i.e. both Atlantic and Gulf Coast landfalling hurricanes), the 

RMW is also positively correlated with latitude, and negatively correlated with the Δp

2

. Using 

only landfall values of RMW the following statistical models best define the relationship between 

RMW, Δp and latitude. 

1. Gulf of Mexico landfalling hurricanes: 

ln(RMW) = 3.558 σlnRMW = 0.457 (8a) 

2. Atlantic Coast landfalling hurricanes: 

ln( ) 2.556 5.963 10 0.0458ψ 5 2 = − × Δ + − RMW p ; r

2

=0.336, σlnRMW = 0.456 (8b) 

3. Gulf and Atlantic Coast landfalling hurricanes: 

ln( ) 2.377 4.825 10 0.0483ψ 5 2 = − × Δ + − RMW p ; r

2

=0.203, σlnRMW = 0.457 (8c) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-118

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 11. RMW for landfalling storms along the Gulf and Atlantic Coasts of the US 

The ability of the RMW models developed using the flight level and H*Wind data (primarily 

open ocean data) to model the landfalling hurricane RMW was tested by computing the mean 

errors (in log space) and the resulting standard deviations and r

2

 values using the landfall RMW

data and the flight level/H*Wind derived RMW models. The error, μlnRMW, is defined as model 

RMW minus observed RMW, thus a mean positive error indicates the model overestimates the 

size of the landfalling hurricanes. The comparisons yield the following findings: 

1. Gulf of Mexico landfalling hurricanes with GoM RMW model: 

μ ln RMW = 0.032 ; r

2

=-0.008 σlnRMW = 0.459 

2. Atlantic Coast landfalling hurricanes with the all hurricane RMW model: 

μ ln RMW = 0.058 ; r

2

=0.356 σlnRMW = 0.450 

3. Gulf and Atlantic Coast landfalling hurricanes with the all hurricane RMW model for the 

Atlantic Coast and GoM RMW model for the Gulf Coast: 

μ ln RMW = 0.043; r

2

=0.219 σlnRMW = 0.453 

0

25

50

75

100

125

150

20 40 60 80 100 120 140

Central Pressure Difference (mbar)

Radius to Maximum Winds (km)

Gulf of Mexico Landfalling Storms

Atlantic Coast Landfalling Storms
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-119

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

A comparison of the model errors noted above to those resulting from the statistical analyses 

of the landfalling storms alone indicates that the models derived from the flight level and 

H*Wind data can be used to define the characteristics of landfalling hurricanes. In the case of 

landfalling Gulf of Mexico hurricanes, the use of the GoM RMW model which contains the 

negative correlation between RMW and Δp

2

, is not statistically significantly different from the 

uncorrelated RMW-Δp relationship derived from the landfalling hurricanes alone. This 

observation suggests that there are an insufficient number of landfalling intense storms in the 

historical data to discern such a relationship. 

Statistical Model for Holland’s Parameter (B) 

The B values computed as discussed above were found to be correlated to the radius to 

maximum winds, central pressure difference, latitude and sea surface temperature. Only points 

associated with central pressures of less than 980 mbar are included in the analysis. Figure 12 

presents the variation of B as separate linear functions of the RMW, Δp, latitude (ψ) and the mean 

sea surface temperature Ts. It is clear from the data presented in Figure 12 that B decreases with 

increasing RMW and increasing latitude. A weak positive correlation of B with Δp is seen as is a 

weak positive correlation with sea surface temperature.

Figure 12. Relationships between the Holland B parameter, latitude, RMW, Δp, and Ts

y = -0.022x + 1.838

R2

 = 0.227

0.00

0.50

1.00

1.50

2.00

2.50

10 20 30 40 50

Latitude (Degrees N)

Holland B Parameter

y = -0.007x + 1.587

R2

 = 0.298

0.00

0.50

1.00

1.50

2.00

2.50

10 20 30 40 50 60 70 80 90 100

RMW (km)

Holland B Parameter

y = 0.003x + 1.113

R2

 = 0.032

0.00

0.50

1.00

1.50

2.00

2.50

20 40 60 80 100 120 140

Central Pressure Difference (mb)

Holland B Parameter

y = 0.059x - 0.391

R2

 = 0.145

0.00

0.50

1.00

1.50

2.00

2.50

10 15 20 25 30

Sea Surface Temperature (Degrees C)

Holland B Parameter
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-120

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

In order to incorporate the effects of RMW, Δp, latitude (ψ) and Ts into a single model, new 

non-dimensional variable, A, was developed defined as: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⋅

Δ ⋅ +

⋅ =

P e

p R T

RMW f A

c

d s

c

2 ln 1

 (9) 

The numerator of A is the product of the RMW (in meters) and the Coriolis force, defined as 

2Ωsinφ and represents the contribution to angular velocity associated with the coriolis force. The 

denominator of A is an estimate of the maximum potential intensity of a hurricane. From 

Emanuel (1988), the maximum wind speed in a tropical cyclone is: 

⎥

⎦

⎤ ⎢

⎣

⎡ =

c

d s p

p V R T max

max 2 ln (10) 

where Vmax is the maximum wind speed, Rd is the gas constant for dry air, pmax is the pressure at 

r=RMW, Ts is the sea surface temperature in degrees K and pc is the pressure at the storm center. 

Using Holland’s Equation it can be shown that 

p e

p

p

p

c c

Δ = 1+ max (11) 

Hence, both the numerator and denominator of A have the units of velocity, and thus A, is 

non-dimensional. Modeling B as a function of the square root of A yields a linear model 

(Figure 13) with B negatively correlated with A and has an r

2

 of 0.34, with a the standard 

deviation of the error equal to 0.225. The relationship between B and A is expressed as: 

B = 1.732 − 2.237 A ; r

2

=0.336, σB = 0.225 (12) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-121

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Figure 13. Relationship between the Holland B parameter dimensionless parameter, A. 

In order to determine if the relationship between B and A is valid for intense storms, the point 

values of B and the model values of B were plotted as a function of RMW for strong hurricanes 

(i.e. storms with a central pressure of < 930 mbar) as shown in Figure 14. The data presented in 

Figure 14 indicates that in the case of strong storms with large RMW (RMW > 40 km) the 

relationship between B and A described earlier breaks down, with the true values of B being less 

than those predicted by the model. Although only two storms with large RMW and low central 

pressures exist in the data analyzed (Hurricane Katrina in the Gulf of Mexico and Hurricane 

Floyd in the Atlantic), the data indicate that the likelihood of a storm with a central pressure less 

than 930 mbar, a RMW greater than 40 km, combined with a B value greater than about 1.1 is 

remote. The mean value of B for these large, strong hurricanes is 1.01, and the estimated 

standard deviation is 0.082. In cases where these strong storms are simulated, B should be 

constrained to lie within the range of 0.85 to 1.18 (i.e. mean ±2σ). 

y = -2.237x + 1.732

R2

 = 0.336

0.00

0.50

1.00

1.50

2.00

2.50

0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4

SQRT(A)

Holland B Parameter
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Figure 14. Holland B parameter vs. RMW for storms with central pressure < 930 mbar 

As in the case of the analysis of Gulf of Mexico (GoM) hurricanes with respect to the 

behavior of RMW with Δp and latitude, B values for all hurricanes within the Gulf of Mexico 

were extracted and analyzed alone. Unlike the results seen for the RMW where the GoM 

hurricanes were found to be smaller than the other hurricanes, the variation of B with A for the 

GoM hurricanes is essentially identical to that seen in the all hurricane case. Figure 15 presents 

the individual B values for the GoM and Atlantic hurricanes along with the model predicted 

mean values of B where it is clearly evident that there is, for practical purposes, no difference in 

the variation of B with A between the two regions. 

Central Pressure < 930 mb

0.00

0.50

1.00

1.50

2.00

2.50

0 10 20 30 40 50 60

RMW (km)

Holland 

B Parameter

Data

Model Mean

Model Mean +/- 2 Std. Dev.

Hurricane Katrina

Hurricane Floyd

Mean B for Large Storms

Mean B +/- 2 Std. Dev.for Large Storms
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Figure 15. Relationship between the Holland B parameter and the dimensionless parameter, A,

comparing the all hurricane data with the GoM hurricane data. 

Note that two simpler, but less elegant models, relating B with RMW and latitude were 

examined modeling B as a function of fcRMW and B as a function of C , where C is defined as: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⋅

Δ ⋅ +

⋅ =

P e

p T

RMW C

c

s ln 1

sinψ (13) 

Where in Equation (13) Ts is expressed in degrees C rather than degrees K. 

The regression model relating B to C is given in the form: 

B = 1.756 − 0.194 C ; r

2

=0.368, σB = 0.220 (14) 

The regression model relating B to fcRMW is given in the form: 

B = 1.793 − 0.326 f cRMW ; r

2

=0.357, σB = 0.221 (15) 

0.00

0.20

0.40

0.60

0.80

1.00

1.20

1.40

1.60

1.80

2.00

0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4

SQRT(A)

Holland B Parameter

Atlantic Hurricanes

Gulf of Mexico Hurricanes

Gulf of Mexico Mean B Model

All Hurricanes Mean B Model
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Both of these models yield marginally improved r

2

 values than does the model relating B to 

the non-dimensional relative intensity parameter, but have the small disadvantage in that the 

independent variable is not non-dimensional. The limitations of these models when applied to 

large intense storms are the same as those evident in the case of the non-dimensional model. The 

reduction the r

2

 value seen when changing the independent variable in the non-dimensional 

parameter given in Equation (9) to the dimensional parameter given in Equation (13) is due 

solely to the conversion of the sea surface temperature from degrees C (Equation 13) to degrees 

K (Equation 9). For practical purpose, any of the three linear regression models given in 

Equations 12, 14 or 15 can be used to model the Holland B parameter, with Equation 15 

requiring the least computational effort. 

Comparisons of Flight Level B Values with Landfall Analysis B Values 

Figure 16 presents a comparison of the Holland B parameters derived from the flight level 

data to those used in the wind field model described in Vickery, et al. (2007) used for estimating 

the wind speeds associated with land falling storms. Overall, the comparisons indicate that the 

two values of B are similar, with the B values used within the windfield model in post storm 

analyses being slightly lower that those derived from the flight level data. The largest difference 

between the two estimates of the Holland B parameter occurs in the case of Hurricane Erin for 

the landfall along the Florida Panhandle. 

Figure 17 presents the same information as Figure 16, but is limited to hurricanes with 

central pressures of 964 mbar or less. The comparisons indicate that the B values used within the 

hurricane wind field model to match the surface observations of wind speeds and pressures is 

about 7% less than those derived from the flight level data. This difference could be due to either 

changes in the characteristics of the pressure field between the 700 mbar level and the surface, or 

biases in the windfield model. Note that a 7% reduction in B corresponds to approximately a 

3.5% reduction in the maximum modeled wind speed. However, comparisons of modeled and 

observed peak gust wind speeds and time series of the surface level pressures suggests that there 

is no such bias in the wind model. 
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Figure 16. Comparison of Holland B parameters derived from flight level data to those derived using a 

post landfall windfield analysis. 

Georges 2 Opal

Frederic

Bonnie

Floyd

Fran

Georges 1

Bertha Erin Elena

Andrew

y = 0.971x

R2

 = 0.728

0.0

0.5

1.0

1.5

2.0

0.0 0.5 1.0 1.5 2.0

Flight Level B

B Used in Wind Modeling
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Figure 17. Comparison of Holland B parameters derived from flight level data to those derived using a post 

landfall windfield analysis for hurricanes with central pressures <= 964 mbar. 

Summary 

The Holland pressure profile parameter, B, was found to decrease with increasing latitude 

and increase with decreasing RMW. A weak positive correlation between B and both Δp and sea 

surface temperature was also observed. The effect of all four of these parameters was accounted 

for by defining a new non-dimensional parameter, A, defined by Equation 12, however; a two 

parameter model (with dimensions) relating B to the RMW and the coriolis parameter is an 

equally good predictor of B.

The limited data for large (as defined by RMW) hurricanes, having low central pressures 

(pc<930 mbar) indicates that B has an upper limit of ~1.2. The relationship between B and A was 

found to be the same in the Atlantic Basin and in the Gulf of Mexico. 

A qualitative examination of the characteristics of intense hurricanes making landfall along 

the coast of the Gulf of Mexico (excluding Southwest Florida) suggests that these hurricanes 

Georges 2 Opal

Frederic

Bonnie

Floyd

Fran

Elena

Andrew y = 0.933x

R2

 = 0.917

0.0

0.5

1.0

1.5

2.0

0.0 0.5 1.0 1.5 2.0

Flight Level B

B Used in Wind Modeling
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weaken in the last 6 – 24 hours, with this weakening characterized by an increase in the central 

pressure, and increase in the radius to maximum winds and a decrease in the Holland B

parameter. The reason for this weakening is beyond the scope of this investigation. 

The few cases where flight level data were available up to the time a hurricane makes 

landfall indicates that in most cases, B, tends to decrease as the hurricane approaches land. 

Recognizing that the data set is limited, this observation suggests that using the statistical model 

for B derived using open ocean (or open Gulf) data may result in an overestimate of B for 

landfalling storms. This potential overestimate of the magnitude of the Holland B parameter 

along the Gulf Coast associated with the use of a statistical model developed using open water 

hurricane data may be further exaggerated because of the decrease in the Holland B parameter 

just before landfall observed in the limited number of landfalling cases examined. 
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Appendix F of R2007 

Integration Method 

There are several ways that one could approach the integration of the storm probabilities 

within the JPM. Each of these can be justified under a particular set of assumptions. Three of 

these will be discussed here; but before we proceed to examine these, a brief overview of the 

integration procedure is in order. 

The integration is performed by summing discretized probabilities to approximate equation 6 

in the main text of this white paper (with ε now included within the same probability function 

as the other five parameters), i.e. 

F1. 

There are three primary aspects of this equation that contribute to the accuracy of 

approximations to F( ) η : 

1. The accuracy in the specification of max ( , , , ,,) p p fl cR v x θ ε ; 

2. The accuracy in the numerical simulations, max ( , , , ,) p fl Ψ cR v x θ ; and 

3. The influence of the discretization size on the approximation. 

The first two of these are fairly intuitive, but the third is often more difficult to explain and 

quantify, since the size of the discrete increments required for a given accuracy is dependent on 

the behavior of the probability function itself. We shall now proceed to examine three 

approaches to the integration. 

1. The Case of Simple Models Combined with Many Simulations 

One approach to estimating coastal hazards would be to utilize simple models with many, 

many simulations. In this context, one could discretize the each parameter in the hurricane 

parameterization considered here (cp, Rmax, νf, θl, x) and even add some additional probabilistic 

attributes to the storms (for example: variations in the Holland B parameter, variable storm 

decay during approach to the coast, suite of variations in offshore storm tracks associated with 

max max max

max

( ) .... ( , , , , , ) [ ( , , , , ) ] ( , , , , , )

where

F( ) is the CDF for surge levels ( )

( , , , , , ) is the probability density function for the multivariate set 

p fl p fl p fl

p fl

F pc R v x H c R v x c R v x

pc R v x

η θ ε θ ηε θ ε

η η

θ ε

= Ψ −+ Δ ∑∑ ∑

max

max

of parameters;

H[z] is the Heaviside function (=1 if z 0, = 0, otherwise);

( , , , , ) is the result of the numerical simulation for a particular parameter combination;

( , , , , , ) is the

p fl

p fl

cR v x

cR v x

θ

θ ε

≥

Ψ

Δ increment of parameter space encapsulated within a discritization.
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each landfalling storm, etc.). Given n parameters with m categories used to represent each 

parameter, nm simulations would be required in this approach. For example, for the fivedimensional representation, with 5 categories in each parameter, 3125 simulations would be 

required to populate the JPM integral. 

This approach would produce a very accurate discretization of the probability function; 

however, inaccuracies in the simulations do not vanish by simply increasing the number of 

simulations. For example, if simplified models predict values that are 30% too high or too low at 

a given point along the coast for the portion of the parameter space that contains the primary 

contribution to surges at the 1 F(η) = 0.01 level (the range of the nominal “100-year” surge 

event), no increase in the number of simulations will converge to the correct answer for this 

range, nor will the addition of uncertainty make the error vanish. 

For the case of wind fields, Vickery et al. (2002) have shown that a parametric wind field 

approach, which includes physical effects captured within an accurate planetary boundary layer 

model, can provide reasonable representations of wind speeds in coastal areas. This method also 

allows many simulations to be run and, therefore, represents a good application of this type of 

approach. Coastal surge generation requires that wave fields, direct wind-driven surges, and the 

interaction between these two driving mechanisms be modeled correctly. Parametric wave 

models for hurricanes have been known to be quite inaccurate for a number of years and are not 

used by any major wave modeling group in the world today, even in offshore areas. In nearshore 

areas, refraction, wave breaking, sheltering and other physical mechanisms cannot be captured 

by parametric models. Also, surge models such as ADCIRC include much improved capabilities 

for depicting critical shallow-water physics and for representing small-scale coastal features. 

Because of this, the methodology described in the main text, which relies on state of the art 

models to ensure unbiased estimates of coastal surges is recommended; however, it is practical 

limitations related to computer run time (even on the largest supercomputers available today) 

presently preclude the execution of multiple thousands of runs using this complete prediction 

system. 

2. The Case of Optimized Category Definition 

Given that a priori information is available for the variation and co-variation of the variables 

used in the probability function, as well as for the sensitivity of coastal surges to variations in 

each of the variables, it is possible to construct an optimized set of discrete samples to simulate 

surges from the overall multivariate distribution. This approach is presently under development 

by Toro for applications within FEMA Region 4. As shown in Figure 7c in the main text, Toro’s 

independent check on the pressure distribution used here showed very good agreement. 

Similarly, preliminary comparisons between the two statistical methods appear to show good 

agreement for JPM applications. 

3. The Case of Structured Interpolation on a Response Surface 

Numerical studies using ADCIRC have shown that coastal surge response is very dependent 

on pressure differential (peripheral pressure minus central pressure), storm size (Rmax), and storm 
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location relative to a site, as discussed in Appendix D. Storm surge is less sensitive to forward 

storm speed and angle of the storm relative to the coast. Figure F1 shows the characteristic 

variation of surge elevations at coastal stations as a function of variations in pressure differential 

(p0 – cp), based on SLOSH tests along the coast of Mississippi. As noted in Appendix B, the 

maximum wind speed in a slowly varying, stationary hurricane can be approximated as 

1/2

1/2

max

a

( )

where

 is the Holland B parameter;

 is the density of air;

 is natural logarithm base (= 2.718...); and

 is the pressure differential between the storm periphery and its cen

a

B U P

e

B

e

P

ρ

ρ

⎛ ⎞ = Δ ⎜ ⎟ ⎝ ⎠

Δ ter.

Since surges tend to be proportional to the wind stress, which for a capped coefficient of drag 

is proportional to the wind speed squared, it follows that the linear relationship between ΔP and 

surge is consistent with our theoretical expectation for this relationship. 

Since a major portion of the surge response to hurricanes is captured by the variation of ΔP

and Rp, the integration method selected for application here is based on the estimation of ΔP – Rp

planes within the 5 dimensional parameter space used in the JPM. Thus, for a fixed value of 

storm landfall location (x), storm track angle relative to the coast (θl), and storm speed (νf), we 

can define a response function, 

max (, ) ( , ,, ) kmn p η x y PR xy = Δ φ F2. 

where φkmn is the surge response function and the subscripts “k, m and n” denote a specific track 

angle, storm speed, and landfall location, respectively. This notation reflects the fact that this 

response function must be defined for each spatial (x, y) point in the computations. Figure F2 

shows an example of such a response function. As expected, the surge values increase essentially 

linearly with increasing pressure differential and also increase with increasing values of Rp. 

Figures F3 and F4 show the characteristic variations of coastal surges as a function of storm 

angle relative to the coast (θl) and forward storm speed (νf), respectively. As can be seen here, 

the variations tend to be quite smooth with either linear or slightly curved slopes in these figures. 

The JPM integration method employed here makes use of the “smoothness” of these functions to 

interpolate between discretized storm parameters. 

An advantage of the approach used here is that the surge response is characterized with some 

level of detail in the important ΔP – Rp plane. To drive the relevance of this point home, let us 

examine the effects of discretization on a probability integral along the “pressure-differential” 
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axis. Given the linear variation of surge elevation with pressure differential, surge elevation can 

been written as 

η = Δλ p F3. 

where λ is a site-specific constant, and hence 

F Fp () ( ) η = Δλ F4. 

If we assume that the pressure differential probability follows a Gumbel distribution, we can 

calculate this probability directly. Figure F5 shows a comparison of the CDF’s for three different 

approximations to the return period as a function of surge level, given the value of λ = 0.2 in 

equation F4 (just for an example here). As can be seen in this figure, for a fixed return period 

deviations between the 3-category approximation (categories with a width of 30-millibars) and 

the continuous function can be as large as ±3.0 feet. For the 3-millibar categories, the deviations 

are, as expected, only about ±0.3 feet. Although this pattern will be smoothed and obscured by 

the addition of many categories of storms, it is clear that the smaller categories provide an 

improved representation for the probabilities. Also, due to the physical basis for the pressure 

differential scaling used here, we can extrapolate to larger storms than those actually utilized in 

the storms. This will provide a somewhat conservative estimate if there is substantial levee 

overtopping; but since the level of extrapolation used here is only about 18% (ΔP = 113 mb to 

ΔP = 133 mb), this should not present a serious problem. 

In this approach, we do not treat the set of 9 storms simulated for a given track, forward 

storm speed, and track angle as a discrete set of storms each with its own associated probability 

increment, since this would give only a relatively crude representation of the actual probability 

structure. Instead, we interpolate between simulated values and extrapolate over relatively short 

distances in Δp and (Rmax). For the integration used in the New Orleans area, the 3 Δp and 3Rmax

values were interpolated to increments of 1-mile in Rmax and 3 millibars in Δp, over the range of 

960 mb to 882 mb for offshore pressures (i.e. before they begin to decay) and 5 nm to 40 nm for 

Rmax values. This provides a very smooth response surface for the primary storm tracks in this 

area (the so-called RICK-fan tracks). For the ±45-degree tracks, the 2 values of Δp and Rmax are 

likewise used to develop a finely discretized (1 nm by 3 mb) set of values over the Δp-Rmax

plane. At this point, we have finely discretized categorizations within the Δp-Rmax plane for the 

five primary (RICK-fan) tracks and the eight additional ±45-degree tracks. Thus, we have a set 

of response functions [φkmn (ΔP, Rp, x, y)] that are quite well defined for the central storm speed 

(11 knots) for these tracks. 

Since the effect of variations in storm speed is fairly small and tends to have fairly linear 

slopes that are roughly independent of (Rmax) and Δp, only a small number of storms can be used 

to modify the functional form of [φkmn (ΔP, Rp, x, y)] as a function of forward storm speed. This 

retains the overall structure of the response function for the alternative speeds, rather than 

assigning only a single surge value to all the responses for that storm speed. For the New Orleans 

area, two storms (with different combinations of Δp and Rmax) were used to estimate the impact 

of a slower storm speed on the surge response function [φkmn (ΔP, Rp, x, y)] compared to the 
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primary (central) speed category (11 knots) for the RICK-fan set of tracks. Single storms were 

also used to estimate the impact of varying the forward speed from 11 knots to 6 knots, along the 

±45-degree tracks and to estimate the effects of varying the forward speed from 11 knots to 

17 knots for all storms. In the approach used here, the value of [φkmn (ΔP, Rp, x, y)] for different 

speeds is obtained from the relationship 

0 0 ( , ,,) ( , ,,) ( , ,,)

where

The subcript "0" refers to the central speed and angle categories

for a specific landfall location; and 

( , ,,) ( , ,,)

kmn p k m n p kmn p

kmn p

kmn p f

f

PR xy PR xy PR xy

PR xy PR xy v

v

φ φ

φ δ

Δ = Δ ΨΔ

∂ Δ

ΨΔ =

∂

For the cases in which a single storm is used to infer the variation with forward speed Ψkmn

(ΔP, Rp, x, y) reduces to a constant. 

This now provides a suitable set of interpolated, finely-discretized values of surge heights as 

a function of Δp and Rmax for all forward storm speeds, storm angles and storm tracks and the 

summation in Equation F1 can be rewritten as 

One concern addressed briefly in Appendix D is the sensitivity of the probability estimates to 

track spacing. The sufficiency of the spacing used here can be investigated by comparing results 

from the runs on the set of tracks that fell between the 5 major RICK-fan tracks to results of 

interpolations based on only the 5 initial tracks. Figures F6-F9 show the results of these 

comparisons. In general little or no bias is introduced into the probability integration, even if 

only the information from the primary tracks is used. Only at Track 4a, where the sites are 

switching from the “right-hand” (onshore winds) side of the storm to the “left-hand” (offshore 

winds) side of the storm in the are east of the Mississippi River does the variability exceed 10% 

of the surge values for surges greater than 10 feet. 

max max max

max

( ) .... ( , , , , , ) [ ( , , , , ) ] ( , , , , , )

where

F( ) is the CDF for surge levels ( )

( , , , , , ) is the probability density function fo

j kl j kl j kl

j kl

i f l mn i f l m i f l mn

i f l mn

F p pR v x H pR v x pR v x

p pR v x

η θ ε θ ηε θ ε

η η

θ ε

= Δ ΛΔ − + ΔΔ

Δ

∑∑ ∑

max

r the multivariate set of parameters;

H[z] is the Heaviside function (=1 if z 0, = 0, otherwise);

( , , , , ) is the interpolated value between simulations with a particular parameter combin j kl i f lm pR v x θ

≥

Λ Δ

max

ation;

( , , , , , ) is the increment of parameter space encapsulated within a discretized category. j kl i f l mn Δ ΔpR v x θ ε
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-133

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Figure F1. Surge levels (feet) at coastal stations (Station Numbers denoted on right hand side of chart) 

along Mississippi coast as a function of pressure differential (mb). 

Figure F2. Response surface showing integerized surge values in feet as a function of row from top to 

bottom (ΔP from 53 to 133 mb in increments of 3 mb) and column from left to right (Rp from 

1 nm to 40 nm in increments of 1 nm). 

Coastal Stations

Varying Central Pressure

0

5

10

15

20

25

20 30 40 50 60 70 80 90 100 110 120

Central Pressure (millibars)

Storm Surge (ft)

15 17

18 20

26 27

29 31

33 34

35 37

38 40

41 44

47 50

51 52

57 63

64 66

72 73

76 82

83 84

95 107

108 119

120 131

132 133

140 141

145 147

Pressure Differential (mb) 148 149
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Figure F3. Surge levels (feet) at coastal stations (Station Numbers denoted on right hand side of chart) 

along Mississippi coast as a function of angle of storm approach to land. 

Figure F4. Surge levels (feet) at coastal stations (Station Numbers denoted on right hand side of chart) 

along Mississippi coast as a function of forward speed of storm (mph). 

Coastal Stations

Varying Angle on Storm Approach

0

2

4

6

8

10

12

14

16

18

20

-45 -30 -15 0 15 30 45

Angle of Approach (deg)

Storm Surge (ft)

15 17

18 20

26 27

29 31

33 34

35 37

38 40

41 44

47 50

51 52

57 63

64 66

72 73

76 82

83 84

95 107

108 119

120 131

132 133

140 141

145 147

148 149

Coastal Stations

Varying Forward Speed

0

2

4

6

8

10

12

14

16

18

20

4 6 8 10 12 14 16 18 20

Forward Speed (mph)

Storm Surge (ft)

15 17

18 20

26 27

29 31

33 34

35 37

38 40

41 44

47 50

51 52

57 63

64 66

72 73

76 82

83 84

95 107

108 119

120 131

132 133

140 141

145 147

148 149
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Figure F5. Comparison of Return Periods for surges estimated from 3-millibar and 30-millibar categories 

compared to a continuous function. 
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Figure F6. Maximum surges produced along a straight east-west coast by storms approaching the coast at 

variable angles compared to storms approaching perpendicular to the coast (tracking due 

north). 

So = 1:10000, Vf = 10 kt, Track Variation

0.00

1.00

2.00

3.00

4.00

5.00

6.00

7.00

0.00 1.00 2.00 3.00 4.00 5.00 6.00 7.00

Surge, Track Due North (m)

Surge, Track at Angle Specified in Legend (m)

15

30

45

-15

-30

-45

match

-5% match

+5% match
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-8-137

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Figure F7. Comparison of results from Track 1a (midway between Tracks 1 and 2) to interpolated values 

using information from Tracks 1 and 2 for a set of points spread throughout the entire New 

Orleans region. 
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Figure F8. Comparison of results from Track 2a (midway between Tracks 2 and 3) to interpolated values 

using information from Tracks 2 and 3 for a set of points spread throughout the entire New 

Orleans region. 
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Figure F9. Comparison of results from Track 3a (midway between Tracks 3 and 4) to interpolated values 

using information from Tracks 3 and 4 for a set of points spread throughout the entire New 

Orleans region. 
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Figure F9. Comparison of results from Track 4a (midway between Tracks 4 and 5) to interpolated values 

using information from Tracks 4 and 5 for a set of points spread throughout the entire New 

Orleans region. 
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Appendix G of R2007 

Estimation of Confidence Bands for Surge Estimates 

Three main types of uncertainty with respect to the estimation of extremes are relevant to 

understanding hurricane hazards along coasts. First, there is uncertainty that the actual sample of 

storms is representative of the “true” climatology today. Second, there is uncertainty in the 

events within future intervals of time, even if the “true” climatology is known exactly. And, 

third, there is uncertainty that some non-stationary process (sea level rise, subsidence, climate 

change, new development patterns, man-made alterations to the coasts, marsh degradation, etc.) 

will affect future hazards. The first of these has traditionally been addressed via sampling theory. 

The second can be addressed via re-sampling or “bootstrap” methods. And, the third must be 

estimated from ancillary information, often not contained within the initial hazard estimates 

themselves. 

The first type of uncertainty listed above pertains to what used to be termed confidence 

bands (or control curves) for estimates of extremes. It cannot be estimated using re-sampling 

techniques, since these techniques use the initial sample as the basis for their re-sampling and 

implicitly assume that the initial sample represents the actual population characteristics. Thus, 

some parametric method must be used to obtain this information. There are many classes of 

distributions which can be used to fit the data. Since we are only using the parametric fits to 

estimate uncertainty and not to replace the non-parametric estimates obtained from the JPM, we 

are somewhat free to use any distribution for which the sampling uncertainty is known. 

Gringorten (1962, 1963) has shown that the expected root-mean-square (rms) error of an 

estimated return period in a two-parameter Fisher-Tippett Type I (Gumbel, 1959) distribution is 

given by 

G1. 

2

T

1.1000 1.1396 1

where

 is the distribution standard deviation;

 is the rms error at return period, T;

N is the number of samples used to estimate the distribution parameters; and

y is the reduce

T

y y

N

σ σ

σ

σ

+ + =

0 1

0 1

d Gumbel variate given by

 ( ) / ;

 is the variate of interest (surge level in this case); and

 and are the parameters of the Gumbel distribution.

y aa

a a

η

η

= −

The reduced variate and return period are related by 
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ln ln

1

T

y T

⎡ ⎤ ⎛ ⎞ = − ⎢ ⎥ ⎜ ⎟ ⎣ ⎦ ⎝ ⎠ − G2. 

which for T >7 approaches an exponential form given by 

1

2

y T e ⎛ ⎞ ⎜ ⎟ − → ⎝ ⎠ G3. 

Equation G1 shows that the rms error at a fixed return period is related to the distribution 

standard deviation and the square root of a nondimensional factor involving the ratio of different 

powers of y (y

2

, y

1

, and y

0

) to the number of samples used to define the parameters. By the 

method of moments, the Gumbel parameters can be show to be given by 

01 1

6 

where is Euler's constant (= 0.57721...) and is the distribution mean.

aa a γμ σ

π

γ μ

=− = G4. 

Thus, the distribution standard deviation is related to the slope of the line represented by 

equation G3. 

Although equation G1 was initially derived for applications to annual maxima, it can be 

adapted to any time interval for data sampling in a straightforward manner. For the case of 

hurricanes, the average interval between storms (the inverse of the Poisson frequency used in the 

compound Gumbel-Poisson distribution) can be used to transform equation G1 into the form 

2

T

1.1000 1.1396 1

where

 is the distribution standard deviation;

 is the rms error at return period, T 

ˆ ˆ (T/T, where T is the average years between hurricanes); and

N is the number of samp

T

y y

N

σ σ

σ

σ

′ ′ + + ′ = ′

′ ′

′ ˆ les used to estimate the distribution parameters (N/T).

 G5. 

Since the form of equation G3 is logarithmic, the slope is not affected by a multiplicative 

factor, and thus, the distribution standard deviation remains the same. N' in equation G5 can be 

estimated from the equivalent total number of years in the sample divided by Tˆ . The total 

number of years for this case is 65 (1941-2005, inclusive) times a factor, Z, which relates the 

spatial area covered by the sample used to the spatial extent of a hurricane surge. For relatively 

intense storms capable of producing surges that are exceeded only every 100 years or more, the 

along-coast extent of very high surges at least 60% of the peak value is about 60 nm for a storm 
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with a 20-nm radius to maximum winds (se Figure D4 in Appendix D). The parameter estimation 

used to derive the values shown in Figure 5 of the main text covered ±3.5 degrees longitude 

along 29.5 north latitude. The value of Z is given by 

Distance along coast

Width of a single sample

Z = G6. 

which in this case is 365.5 nm divided by 60 nm, or approximately 6.1. Thus, the effective 

number of years is 396. 
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Appendix 8-3 

Rainfall Analysis 

Rainfall Intensity 

Rainfall is an additional mechanism that contributes to the inundation of the sub-basins. 

While rainfall is not of primary concern for the hurricane protection system, it is a contributor to 

the frequency of low-level flood losses. Hence it was decided that a relatively coarse model of 

hurricane-induced rainfall would suffice. 

Prior to NASA’s Tropical Rainfall Measuring Mission (TRMM) (Simpson et al., 1988), 

information on hurricane rainfall was scanty. The TRMM mission, which started in November 

1997, produced significant rainfall data for tropical storms and hurricanes at a spatial scale of 

about 5 km in various tropical regions, including the Atlantic basin. These rainfall products have 

been analyzed statistically by Lonfat et al. (2004) and Chen et al. (2006). The model proposed 

below is based primarily on those two studies and on discussions with Dr. Shuyi Chen at the 

University of Miami. 

Mean Rainfall Intensity 

Hurricane rainfall intensity I (mm/hr) varies with distance r from the hurricane center and 

azimuth β relative to the direction of motion. Moreover, the mean intensity field mI (r, β) varies 

with the central pressure deficit ΔP, the radius of maximum winds Rmax, the storm velocity V, 

and the vertical wind shear S (in the above quoted references, S is measured as the difference 

between the horizontal wind fields at the 200 and 850 hPa levels). Finally, rainfall intensity

displays strong fluctuations at different scales around the mean value mI (r, β). 

The azimuthal average of mI (r, β), mI (r), gives the symmetrical component of the mean 

rainfall field. This component has a maximum at a distance from the hurricane center close to 

Rmax and decays in an approximately exponential way at larger distances. This decay is 

contributed by the approximately exponential decay of both the fraction of rainy area and the 

mean rainfall intensity at the rainy locations. The rate of exponential decay mI (r) is inversely 

proportional to the size of the hurricane; hence, in good approximation, it is inversely 

proportional to Rmax. 

The value of mI (r) for r = Rmax increases with increasing ΔP, approximately doubling from a 

Cat2 to a Cat4-5 event. Considering the Cat12 and CAT3-5 results in Lonfat et al. (2004) as 

representative of the Cat1-2 boundary and of Cat4, respectively, assuming linear dependence of 

the mean rainfall intensity at Rmax on Δp, and fitting an exponential decay with distance as 

mentioned above, one obtains 
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( )

( ) max

max

max

0.3

max

1.14 0.12 ,

1.14 0.12 ,

r R I

R

P for r R

m r

P e for r R

⎛ ⎞ − − ⎜ ⎟ ⎝ ⎠

⎧ +Δ ≤ ⎪ = ⎨

⎪ +Δ > ⎩

 (10) 

where mI is in mm/h and ΔP is in mb. 

The asymmetric component of the mean rainfall field, i.e. the way mI (r, β) depends on the 

azimuth β, is affected mainly by the storm velocity V and the vertical wind shear S. This 

influence is complex, as the asymmetric pattern and its strength vary with the absolute and 

relative values of V and S, the relative direction of wind shear and storm motion, the distance r

from the center, and the geographic location. For hurricanes in the Atlantic region, there is a 

general tendency for rainfall to intensify in the front-east quadrant relative to the direction of 

storm motion and de-intensify in the rear-west quadrant. This tendency is especially evident for 

fast-moving storms and away from the hurricane center, reaching about 30-40% of mI (r) for 

r ≈ 3Rmax. The effect is stronger over land than over water. 

Variability of Rainfall Intensity 

For each TRMM observation of each hurricane, Lonfat et al. (2004) extracted the average 

rainfall intensity I rr ( ) , 10 + + at rainy locations inside annular regions of 10 km width. Using 

these values, they found the empirical distribution of I rr ( , 10) + + for different r and different 

storm intensity classes. A consistent result is that I rr ( , 10) + + varies by a factor of about 7 

above and below the median value. The standard deviation of log , 10 (I rr ( )) + + corresponds to 

a factor of about 2-2.5. Hence the variability of this average rainfall intensity is very large. 

In addition, there is variability in the fraction of rainy area. The latter variability is not given 

in Lonfat et al. (2004), but it can be bounded and roughly estimated as follows. The mean 

fraction of rainy area, mF+, is given by Lonfat et al. as a function of r and storm intensity range. 

Given mF+, an upper bound to the variance of F+

 is obtained by assuming that F+ is either 0 (no 

rain in the region) or 1 (it rains everywhere in the region). In this case ( ) 1

F F Var F m m + +

+ ⎡ ⎤ = − ⎣ ⎦ , 

with a coefficient of variation 1 1

F

F

V

m +

+

= − . A more realistic estimate of the coefficient of 

variation is perhaps one half of this theoretical upper bound, or 

VF+ ≈ 0.5 1

mF+

−1 (11) 

For distances r up to 150 km, which are those that contribute the most to intense rainfall, mF+

is around 0.9 irrespective of hurricane intensity and Eq. 11 gives VF+ ≈ 0.17. This coefficient of 
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variation is much smaller than the coefficient of variation of rainfall intensity inside the rainy 

area, which is on the order of 1.0. Therefore, the variability of the rainy area may be neglected. 

Assessment of Rainfall Intensity inside the Sub-basins 

Based on the above considerations, the following simplified model of rainfall inside the 

sub-basins was adopted. First, the mean rainfall contribution from the symmetric component of 

the mean rain field, mI (r) is specified; then the asymmetric component is discussed, and finally 

an assessment of the variability of rainfall around the mean is given. 

Denote by ( ) , k m rt I the temporal variation of mI (r) for hurricane k (in this model, temporal 

variation is due to the variations of ΔP). The contribution of ( , ) k m rt I to the mean rainfall 

intensity in subbasin j is evaluated as ( ( ), ) k m r tt I jk , where rjk(t) is the distance to a 

representative point of subbasin j from the center of hurricane k at time t. The appropriate 

distance was taken to be the instantaneous distance between the eye of the storm and the centroid 

of the subbasin. 

For hurricanes that pass to the right or near the subbasin, one may conservatively neglect the 

azimuthal dependence of the rainfall field. For hurricanes that pass to the left of the sub-basin 

centroid, the asymmetric component has been accounted for by multiplying the above symmetric 

mean rainfall values by 1.5. This factor includes intensification due to land effects. 

Uncertainty may be expressed by a lognormal random variable with mean value 1 and log 

standard deviation 0.69, which corresponds to an uncertainty factor of 2. This random factor 

should be applied to the entire mean rainfall time history. In reality, rainfall intensity inside a 

subbasin would display significant fluctuations in time and space, which locally could far exceed 

a factor of 2. However, the above random factor should adequately reflect uncertainty on the 

total precipitation in a sub-basin during the passage of a hurricane. 
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Appendix 9 

Risk Methodology 

Introduction 

This Appendix describes the risk analysis philosophy and methodology used to evaluate the 

performance of the New Orleans hurricane protection system. Probabilistic risk analysis as 

described by Ayyub (2003), Kumamoto and Henley (1996), and Modarres et al. (1999) was used 

to develop the basic risk analysis methodology of the hurricane protection system. The basic 

elements of the risk analysis methodology are illustrated in the flow chart presented in 

Figure 9-1. The analysis was developed as a series of modules which interface to provide a risk 

model for the New Orleans HPS. An Excel spreadsheet program, Flood Risk Analysis for 

Tropical Storm Environments (FoRTE), was developed to implement the many water volume 

calculations and exceedence values required to determine the risk of inundation for the suite of 

hurricanes investigated. The spreadsheet (FoRTE) is described in Appendix 17. The results of 

the many FoRTE program runs were post processed and modified to include wave runup, 

interflow between sub-basins for the aggregated storm water volumes, and pumping, and to 

adjust the program outputs based on historic experience. The results of the analyses are 

described in Appendix 13. 

In the engineering community, risk is generally defined as the potential that a component or 

system will incur losses from exposure to a hazard or as a result of an uncertain event. Risk is 

quantified as the rate (measured in events per unit time) that lives, economic, environmental, and 

social and cultural losses will occur due to the non-performance of an engineered system or 

component. The non-performance of the system or component can be quantified as the 

probability that specific loads (or demands) exceed respective strengths (or capacities) causing 

the system to fail, and losses if that failure occurs. Risk can be viewed to be a multi-dimensional 

quantity that includes event-occurrence rate (or probability), event-occurrence consequences, 

consequence significance, and the population at risk; however, it is commonly measured as a 

pair of the rate (or probability) of occurrence of an event, and the outcomes or consequences 

associated with the event’s occurrence that account for system weakness, i.e., vulnerabilities. 

Another common representation of risk is in the form of an exceedance rate (or probability) 

function of consequences. In a simplified form, risk is commonly expressed as: 

Risk = Event rate (or probability) × Vulnerability × Consequences of failure 
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This equation not only defines risk but also offers strategies to control or manage risk, i.e., by 

making the system more reliable or by reducing the potential losses resulting from a failure. The 

vulnerability, or probability of failure, part of the equation can be influenced by engineers by 

strengthening of existing structures, increasing reliability or by adding additional protection. 

However, the consequence part is highly dependent upon the actions and decisions made by 

residents, government and local officials, including first-response and evacuation plans and 

practices. In densely populated areas, simply increasing system reliability may not reduce risks 

to acceptable levels and increasing consequences through continued flood plain development can 

offset any risk reductions or cause an increase in risk. 

A reliability analyses is used to model the performance of individual elements and features 

(such as, floodwalls, levees, pumps, levee closures, etc.) located throughout the hurricane 

protection system to the overall performance of the integrated HPS. The reliability of the various 

elements and features considers the varying material properties of the structures and of 

foundation conditions that exist throughout the HPS. The impact of this performance on public 

safety and, social and economic welfare is incorporated into the risk analysis. 

Implementation of risk analysis to the HPS of New Orleans and S.E. Louisiana was 

challenging because it is a complex system of levees, floodwalls and pumping stations, 

constructed over many years by different entities that serve a large geographical region. In 

addition, existing capability to accurately predict and model hurricanes in regions as complex as 

the Mississippi delta was limited. Nonetheless, mathematical modeling of hurricanes and risk 

analysis methodologies have improved greatly in recent years to make them important, viable 

tools for supporting investment decisions as the HPS is restored and improved. In developing the 

risk analysis strategy, the following requirements were identified as key guiding principles: 

• Analytic. The methodology must provide a systematic framework for assessing risk by 

decomposing risk into its basic elements. 

• Transparent. All assumptions and analytical steps are clearly defined. 

• Defensible. Values for each parameter are supported by all available data, including 

knowledge from previous studies and expert opinion. 

• Quantitative. Risk is expressed in meaningful and consistent units (e.g., dollars and 

fatalities) so as to provide a basis for performing tradeoffs and benefit-cost analysis. 

• Probabilistic. The mathematics of probability theory is used for expressing uncertainty 

in all model parameters and assessing the likelihood of alternative scenarios. 

• Consistent. It is consistent with established and accepted practices of probabilistic risk 

assessment (PRA) used in many other fields. 

The quantification of risk also required that the analysis consider uncertainty in both the 

input values and the modeling capabilities. For example, detailed knowledge of the engineering 

parameters that influence the performance of the HPS and of the hurricane characteristics of 

storms expected to impact New Orleans is limited. This includes properties of foundation soils 

underlying the extensive levee and floodwall system, and the frequency with which hurricanes 

will occur in the future. As other examples, Dixon, et al. (2006) provides an overview of 

subsidence and flooding in New Orleans; Dokka (2006) describes the tectonic subsidence in 

coastal Louisiana; and Muir-Wood and Bateman (2005) describe uncertainties and constraints on 
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breaching and their implications for flood loss estimation. Hurricane models can predict winds, 

waves and surges only with limited accuracy, and the reliability models used to predict levee 

performance when subjected to hurricane forces are similarly limited. Hence, the risk profiles of 

hurricane-induced flooding cannot be established with certainty. Risk analysis, therefore, must 

include not just a best estimate of risk, but also an estimate of the uncertainty in that best 

estimate. By identifying the sources of uncertainty in the analysis, measures such as gathering 

additional data can be taken to reduce the uncertainty and improve the risk estimates. 

Several key considerations and limitations of the IPET risk study which should be noted are: 

• Defining the physical features of the system required an accurate inventory of all 

components that provide protection against storm surge and waves. This included cross 

sections, strength parameters of components, transitions between elements, crest 

elevations and foundation conditions along reaches. The characterization of the physical 

features of the protection system was, however, limited by the availability of up-to-date 

information, the resources to conduct detailed field surveys, and the ability to process the 

large amount of information that was changing during the course of the study. 

• The hurricane modeling and reliability analyses required an accurate depiction of the 

elevations of the tops of levees and walls that make up the HPS. This was complicated by 

the different datum used in the area over many years, the lack of up-to-date pre-Katrina 

survey data and the damage caused by Katrina. The risk team utilized the work by other 

IPET teams to define the datum to be used. The datum used for all elevations cited in the 

risk study was NAVD88 2004.65. The risk team used data provided by the New Orleans 

District, Task Force Guardian and others to establish the pre- and post-Katrina crest 

elevations. 

• The pumping system is an important element of the HPS that controls flooding during 

and after rain and tropical storms, but was not designed to handle overtopping and 

breaching during hurricane events. This is also complicated by the human factors that 

affect the operation of the pumping system. For these reasons, several levels of pumping 

performance were investigated to provide a range of potential performance levels. 

• The consequences associated with pre- and post-Katrina flooding are different due to 

changes in population and economic activity. 

• The effectiveness of the protection system depends on human factors as well as 

engineered systems (e.g., timely road and railroad closures, gate operations, and 

functioning of pumping stations). Lessons learned from Katrina and other natural 

disasters were used in modeling the closures. 

• Wave runup, interflow between sub-basins for aggregated storm water volumes, and 

pumping was considered outside of the FoRTE program using a simplified analysis. 

Adjustments to the FoRTE outputs were based on historic data. 
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Key Factors Influencing Risk 

The development of a risk analysis model was facilitated by the preparation of an influence 

diagram. The process of creating an influence diagram helped establish a basic understanding of 

the elements of the hurricane protection system and their relationship to the overall system 

performance during a hurricane event and defined input required for the analysis of 

consequences and risks. 

Figure 9-2 shows the influence diagram for the hurricane protection system and the analysis 

of consequences. There are four parts to the influence diagram: 

• Value nodes (rounded-corner box) 

• Chance nodes (circular areas) 

• Decision nodes (square-corner boxes) 

• Factors and dependencies in the form of arrows. 

The influence diagram was used to develop an event (or probability) tree for the hurricane 

protection system. Figure 9-3 shows an initial probability tree derived from the influence 

diagram in Figure 9-2. The top events across the tree identify the random events whose state 

following the occurrence of the hurricane could contribute to flooding in a protected area. The 

tree begins with the initiating event which is a hurricane that generates a storm surge, winds and 

rainfall in the region. 

Analysis Boundaries 

An important initial step in the analysis is to clearly define the bounds of the study and the 

physical descriptions of the various components of a HPS. These bounds included defining the 

geographic bounds of the study region, the elements of the hurricane protection system, the 

resolution of information and analyses to be performed, and analysis constraints or assumptions 

associated with the risk and reliability analyses. 

Study Region and Hurricane Protection System 

Figure 9-4 identifies the region of southeast Louisiana considered and the major parishes of 

the area protected by the hurricane protection system. The HPS study area is limited to the six 

parishes that make up the metropolitan New Orleans area. 

Physical Description of the HPS 

The HPS is comprised of a variety of sub-systems, structures, and components, which 

include earthen levees, floodwalls, pumping stations, drainage canals, road and railway closures, 

and power supply systems. The system is a combination of low lying tracts surrounded by flood 
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barriers that form drainage basins, which are independently maintained and operated by local 

parishes and levee boards. Detailed physical descriptions for each basin based on current 

conditions are provided by Appendices 2 through 7. Data collected during site inspections by the 

risk team were used to define characteristics of the basins and their interdependence for use in 

the risk model. This was a critical and time consuming step in the development of the risk model 

that has yielded a comprehensive description of the HPS. These descriptions were developed by 

examining available information gathered by IPET including: 

• Design memorandums and supporting documents, 

• Pre- and post-Katrina construction documents, 

• Inspection reports, 

• Katrina damage reports, and 

• Detailed field surveys conducted by the Risk Team to verify the location and 

configurations of the HPS. 

• Comprehensive studies conducted by other IPET teams 

• Information collected by Task Force Guardian during repair of the HPS. 

The information gathered was incorporated into detailed geographic information system 

(GIS) based maps of each basin that included: locations of all features (walls, levees, pumping 

stations, and closure gates), geotechnical information (boring logs, geologic profiles), aerial 

photographs, photos of each feature and elevations of the tops of levees and walls. 

Analysis Assumptions and Constraints 

As part of the process of developing the risk analysis model, it was necessary to identify key 

assumptions and analysis constraints. Constraints refer to events or situations that were not 

modeled or considered explicitly in the analysis. The analysis limitations or constraints of the 

risk model development are summarized by the following: 

• Only modeling procedures that existed prior to Katrina were used. 

• Geographic area was limited to elements of the hurricane protection system in the 

following basins: 

o St. Charles 

o Jefferson (East and West Bank) 

o Orleans (East and West Bank) 

o New Orleans East 

o St. Bernard 

o Plaquemines 

• The risk model does not produce temporal profiles, but rather spatial profiles 

accumulated over the durations of respective storms. 

• The risk model includes assumptions based on the information collected to select the 

parameters used in various major aspects of the hurricane protection system 
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characterization, hurricane simulation, reliability analysis, inundation analysis, and 

consequence analysis. 

• Hazards, and thus consequences, not considered in the risk analysis are: wind damage to 

buildings, fire, civil unrest, indirect economic consequences, effect of a release of 

hazardous materials, and environmental consequences. 

• The performance of the evacuation plan New Orleans was not modeled in the risk 

analysis. 

Hurricane Protection System 

The hurricane protection system (HPS) for the New Orleans metropolitan area shown in 

Figure 9- 4 is sub-divided into basins that follow parish boundaries and sub-basins that define 

the interior drainage characteristics of the basins. Basins and sub-basins are divided into 

sections, or reaches, that have similar cross-sections, material strength parameters and 

foundation conditions. Features such as: pumping stations, road and railway closures, drainage 

structures, etc. within a reach are defined as points within the reach that have the potential for 

allowing water inflow in the event of their failure. The HPS has been discretized for the 

reliability and risk analysis tasks as schematically shown in Figure 9- 5 which shows an example 

of how the HPS was discretized to define the system in the risk model. A complete definition of 

the system is provided in the appendices. The system consists of basins, sub-basins, reaches, 

features, and transitions. The definition of these components of the HPS are based on the 

following considerations: 

• Local jurisdiction, 

• Floodwall type and cross section, 

• Levee type and cross section, 

• Engineering parameters defining structural performance, 

• Soil strength parameters, 

• Foundations parameters, 

Reaches of each basin are uniquely identified using sequential numbers as illustrated in the 

Figure 9-5. The figure also shows the approximate locations of pumping stations for the purpose 

of illustration. 
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Definition of Basins, Sub-basins, Reaches and Features

The hurricane protection system is divided into basins, sub-basins, reaches, features and 

transitions. Table 9-1 illustrates the information structure needed for this definition for selected 

reaches. The definition includes the following basins with their respective numeric identification: 

1. Orleans West Bank (OW) 

2. New Orleans East (NOE) 

3. Orleans (OM) 

4. St. Bernard (SB) 

5. Jefferson East (JE) 

6. Jefferson West (JW) 

7. Plaquemines Area (PL) 

8. St. Charles (SC) 

Reach Descriptions 

The HPS perimeter is discretized into reaches that define sections that have similar physical 

and engineering characteristics. Initially the reaches were defined using the beginning and 

ending stations shown in the design memoranda (DM). The stations were then adjusted based on 

examinations of the subsurface material information to form reaches that were expected to have 

similar performance (reliability). For each reach, the following information, as shown in 

Table 9-1, is required: 

• Reach numeric identification that can be associated with a unique station in hurricane 

simulation 

• Reach length (ft) 

• The reach crest elevation (ft) 

• Reach type, either a levee (L) or a floodwall (W) 

• Reach weir coefficient needed to compute overtopping water volume of either 2.6 for a 

levee or 3.0 for a floodwall (in units of ft and sec) 

• Basin reference that defines the location of the reach in reference to the overall HPS 

• Sub-basin reference that defines where water from overtopping or breaching of the reach 

will collect. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-8

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 9-1. Definition of Reaches 

Reach 

No. 

Length 

(ft) 

Elevation 

(ft) 

Design Water Elevation 

(ft) 

Reach 

Type 

Reach Weir 

Coefficient 

Subbasin 

Reference 

1 5,000 14.00 8.00 Levee 2.6 Basin1-1 

2 10,000 15.00 9.00 Floodwall 3.0 Basin1-2 

3 22,500 16.00 11.00 Levee 2.6 Basin1-3 

4 6,000 14.00 10.00 Floodwall 3.0 Basin1-4 

5 9,000 18.00 13.00 Levee 2.6 Basin1-5 

6 7,000 14.00 8.00 Levee 2.6 Basin2-1 

7 11,000 15.00 9.00 Floodwall 3.0 Basin2-2 

8 7,500 16.00 11.00 Levee 2.6 Basin2-3 

9 500 11.00 8.00 Transition 3.0 Basin1-2 

10 400 12.00 8.00 Transition 2.6 Basin2-2 

Feature Descriptions 

Table 9- 2 illustrates the definitions of features within each reach for selected reaches. For 

each feature, the following information is required: 

• Feature number for unique identification 

• Type of features of drainage structure (D), or closures (i.e., gate G), or transition 

structure (T) 

• Reach reference where the feature is located 

• A reference value for correlated gates for assigning the same probability of closure 

• Width of opening (ft) for water inflow through open gates 

• Bottom elevation (ft) of gates 

• Probability of gate not closed during a hurricane 
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Table 9-2. Definition of Features with Respective Reaches 

Feature No. Reach No. Correlated Features Length (ft) Bottom Elevation (ft) Not-closed Probability 

1 1 1 500 5.00 0.10 

2 1 1 500 5.00 0.15 

3 2 3 400 6.00 0.10 

4 2 3 400 7.00 0.20 

5 2 3 400 5.00 0.10 

6 3 3 600 5.00 0.15 

7 4 7 600 7.00 0.20 

8 4 8 600 6.00 0.10 

9 5 9 500 6.00 0.10 

10 5 9 500 5.00 0.01 

Sources of Information 

The Risk Team collected data from design documents, construction drawings and studies 

conducted by other IPET teams to develop detailed descriptions of the basins. Maps were 

assembled from aerial photos and information was overlayed in GIS files that included: lat/long 

data, geotechnical profiles and boring logs, crest elevations, stationing used to define reaches 

and the locations of critical features such as closure gates and pump stations. The information on 

these maps was confirmed by field surveys of the entire system by members of the Risk Team 

who traveled every mile of the system. Photos, GPS coordinates and notes were taken during 

these surveys to document each feature and reach used in the risk model. In addition to the maps, 

data was compiled for use in the reliability analyses and the risk model. This process has resulted 

in a comprehensive description of the HPS. The basin descriptions are provided in Appendices 2 

thru 7. 

Elevations of Crests 

The elevations of the tops of walls and levees, adjusted to the current datum, of the entire 

New Orleans area HPS were developed for use in the suite of hurricane simulations and the risk 

assessment model calculations of water volumes from overtopping and breaching. Various 

sources for elevations of segments of the HPS existed, some adjusted to current datum, but most 

were not. The 1 ft2

 and 15 ft2 lidar data on the IPET repository have been adjusted to current 

datum and gave about a 99% coverage of the HPS system. The adjusted lidar data gave good 

values for portions of the HPS that had levees that were clear of vegetation. In addition, there 

were numerous field surveys that were available for short portions of the walls, some of which 

been adjusted to the latest datum. 

Using the 1 ft2

 lidar where it was available, cross section profiles were created for lengths of 

approximately 200 to 500 ft along the entire HPS. Where the 1 ft lidar was not available, the 

15 ft2

 lidar was used. For the levees, these elevations were compared to the current expected 

values obtained from various MVN records, Taskforce Guardian, and any available field survey 

information for verification. The location of walls, drainage structures, closures, and gaps were 
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known from the field survey of the entire HPS that was was documented with photos and notes. 

Some walls had adjusted survey information available, but for most walls it was necessary to go 

back to the lidar data and examine the areas by drawing numerous profiles, searching for lidar 

data patterns of “good hits” on wall tops and determining the elevations of the surrounding soil. 

Then, using the photos and notes obtained from the site visits, estimates of the wall elevations 

were made. This same process was used for transition regions. A final comparison to the 

elevations used in the grid developed by the Storm Team for use in the computer program 

ADCIRC was made for consistency. 

Performance of HPS Structures

The performance of the structures providing hurricane protection against potential water 

elevations due to surge and waves was quantified using structural and geotechnical reliability 

models integrated within a larger system description of each drainage basin. The reliability 

models for the HPS components were developed based on design and construction information, 

and on the results of the IPET Performance Team and the Pump Stations Team studies. 

Reliability models were developed and evaluated to determine dominant, or most likely, failure 

modes for each reach defined in a drainage basin. Failure modes, performance functions, basic 

random variables, and computational procedures used to model failure probability are provided 

in Appendix 10, Reliability Methodology. 

The reliability models included uncertainties in structural material properties, geotechnical 

engineering properties, subsurface soil profile conditions, and engineering performance models 

of levees, floodwalls, and transition points. Uncertainties due to spatial and temporal variation, 

and due to limited knowledge are tracked separately in the analysis. The reliability models 

provided a best estimate of the frequency of failure under given loads, along with a measure of 

the uncertainty in that frequency. 

Engineering performance models and calculations were adapted from the Geotechnical 

Design Manuals (GDM) and engineering parameter and model uncertainties were propagated 

through those calculations to obtain approximate fragility curves as a function of water height for 

components of the HPS. These results were calibrated against the analyses of the Performance 

Team, which applied more sophisticated analysis techniques to similar structural and 

geotechnical profiles in the vicinity of failures. Failure modes identified by the Performance 

Team were incorporated into the reliability analyses as those results became available. 

Reliability assessments were performed for individual reaches of the HPS for given water 

elevations. The assessments resulted in fragility curves for each reach by dominant or most 

likely mode of failure. A fragility curve gives the probability of failure, conditional upon an 

event (water elevation in this study), at which a limiting failure state is exceeded. A sample 

fragility curve is shown in Figure 9-6 and the actual curves used in the risk analysis are shown in 

Appendix 10. 
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Hurricane Hazard Analysis 

The hurricane hazard analysis method parameterizes hurricanes using their characteristics at 

landfall. The hazard analysis was conducted by an independent team that included 

representatives from USACE, FEMA, consulting firms and academia. Details of the analysis are 

presented in Appendix 8. The following parameters were considered: 

• Central pressure deficit at landfall, 

• Radius to maximum winds at landfall, 

• Longitudinal landfall location relative to downtown New Orleans, 

• Track of storm motion at landfall, 

• Storm translation speed at landfall, and 

• Holland’s radial pressure profile parameter at landfall (Holland 1980). 

Using parameter values based on historic events, the recurrence rate and the joint probability 

density function of the hurricane parameters were estimated for hurricane events in the New 

Orleans region of interest. The parameters used by the hurricane team to develop the storms they 

provided to the risk team are shown in Table 9-3. Note that frequencies of the storms highlighted 

in yellow in Table 9-3 were not provided; therefore, these hurricanes were not used in the risk 

analysis. 

The selected hurricanes were used as input to the ADCIRC models which used several finite 

element grids for the various conditions of the HPS. The grids corresponded to the condition of 

the HPS before Katrina and after repairs and improvements had been completed following 

Katrina. 

Since the possible combinations of winds, surges and waves would be computationally 

demanding if every combination was run through the ADCIRC models, the number of runs was 

reduced by using a response surface approach. In this approach a relatively small number of 

hurricanes were selected and used to calculate the corresponding surge and wave levels at the 

sites of interest. Then a response surface model was fitted to each response variable (surge or 

wave level at a specific site). Finally, a refined discretization of the parameter space was used 

with the response surface to represent the hurricane hazard. The outcomes of these computations 

were combined surge and effective wave setup elevations at particular locations of interest along 

the hurricane protection system, e.g., representative values at points along the reaches. 

A hydrograph with time-varying surge plus wave setup elevations at each reach was 

produced, based on the ADCIRC analyses, and provided as input to the risk model. Example 

hydrographs for a single reach are show in Figure 9-7. Wave runup elevations were added using 

a simplified approach. 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Risk Quantification 

The quantification of risk associated with a hurricane protection system required establishing 

a performance measure for the HPS. The selected performance measure was the amount of water 

expected to enter protected areas during a particular hurricane. The water entered protected areas 

as a result of one or more of the following cases: 

1. Non-breach events such as overtopping, water entering through closures (i.e., gates) that 

are left open, precipitation, and potential backflow from pumping stations 

2. Breaching events caused by levee or flood wall failure that lead to water inflow into 

protected areas 

3. Rainfall during hurricane events 

The risk quantification framework has, therefore, the objective of estimating water volumes 

and elevations in basins according to these cases. The event tree presented in Figure 9-8 shows 

the quantities of interest in the net water levels (W) column resulting from open closures, 

overtopping, breaching, and operation of pumping stations in non-breach cases. The branches for 

the rainfall volume are shown separately for clarity, but were added to all the other branches 

during calculations. Figure 9-8 shows a total of 12 branches that were evaluated for each 

hurricane. These branches were numbered sequentially as shown in the event tree. The top 

events of the tree are defined in Table 9-3. 

The results for each storm event were evaluated by aggregating the individual results for 

each basin. This provided an estimate of water inundation volumes in each basin along with the 

frequencies of occurrence. These were then converted to elevations using the stage-storage 

curves for each basin thereby yielding elevation – exceedence relationships. The evaluation of 

consequences relative to basin inundation levels were provided by the Consequence Team based 

upon the elevations selected for the 50, 100 and 500 year events. 

Table 9- 3 

Summary of the Event Tree Top Events 

Top Event Description 

Hurricane 

initiating event 

The hurricane initiating event maps the peak flood surge and wave effects with a hurricane rate λ. This event 

can be denoted, hi(x,y), and has a probability of occurrence, P(hi(x,y)) and a rate of occurrence of λP(hi(x,y)). 

Closure structure 

and operations 

(C) 

The closure event models whether the hurricane protection system closures, i.e., gates, have been sealed 

prior to the hurricane. This event depends on a number of factors, as illustrated in the influence diagram. The 

closure structures were grouped by basins in terms of probability of being closed in preparation for the arrival 

of a hurricane. This event can be used to account for variations in local practices and effectiveness relating to 

closures and their operations. 

Precipitation 

inflow (Q) 

The precipitation event models the rainfall that occurs during a hurricane event. The precipitation inflow per 

subbasin is treated as a random variable. 

Drainage, 

pumping and 

power (P) 

The drainage event treats pumping in aggregate with drainage effectiveness and power reliability, including 

backflow through pumps. 

Overtopping (O) This event models the failure of the HPS due to overtopping. 

Breach (B) The breach event models the failure of the HPS during the hurricane. This event includes all failures other 

than overtopping. This event is treated using conditional probabilities as provided in Figure 9-8. 
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Risk associated with the hurricane protection system was quantified through a regional 

hurricane rate (λ) and the probability P(C > c) where a consequence measure C exceeds different 

levels c. The loss exceedance probability per event was evaluated as 

( ) ()( | )( | , ) i ji ij

i j

PC c Ph PS h PC c h S >= > ∑∑ (9-1) 

An annual loss exceedance rate was estimated as follows 

( ) ()( | ) ( | , ) i ji ij

i j

λ λ C c Ph PS h PC c h S >= × > ∑∑ Eq. (9-2) 

where P(hi) is the probability of hurricane events of type i, P(Sj|hi) is the probability that the 

system is left in state j from the occurrence of hi, and P(C > c | hi, Sj) is the probability that the 

consequence C exceeds level c under (hi, Sj). Summation was over all hurricane types i and all 

system states j in a suitable discretization. Simulation studies of hurricanes for risk analysis 

required a set of hurricane cases hi and their respective rates of occurrence λi. 

Evaluation of the regional hurricane rate λ and the probability P(hi), the conditional probabilities P(Sj | hi), and the conditional probabilities P(C > c |hi, Sj) was obtained from the 

hurricane model, the HPS risk assessment model, and the consequence model, respectively. 

Water Inflow Volume Models 

The hydrographs and HPS system descriptions and fragilities were used to compute whether 

water entered a basin by levee overtopping or breach, and to determine the resulting water 

elevation (Hps) within the basin. In the case of levee overtopping, Hps within a basin was based 

on the water volume computed using the duration of overtopping. If a breach occurred and the 

invert of the breach was below the final elevation of an adjacent body of water, Hps was set to the 

elevation of that body of water. If the breach invert was above the final elevation of an adjacent 

body of water, Hps was based on a water volume computed using the duration that the surge 

elevation was above the breach invert. The topography, stage-storage curves, and the drainage 

and pumping models for a basin were used to construct such a relationship. The major basins 

were subdivided into sub-basins according to the drainage and pumping characteristics within 

the basin. These subdivisions are show in Figure 9-9. 

Water Volumes from Other Features of the Protection System. The hurricane protection 

system includes features that could allow water volume to enter the protected areas during a 

hurricane. These features include: 

1. Closure structures, i.e., gates, that were left open or failed to close 

2. Local changes in elevations at transitions in the HPS, typically between levees and 

floodwalls 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

These features are identified within each reach and assigned to a subbasin. The water volume 

resulting from failure of closure structures for a given hurricane was computed using the closure 

structure failure probability, width of the closure structure, and the elevation at the bottom of the 

structure. The water volume associated with localized changes in transitions required the change 

in elevation and the lengths over which the elevation varied. 

Table 9-4. Sample Reach Overtopping Volume Results 

Overtopping Volume (ft3) 

For Basin1-1 For Basin1-2 For Basin2-1 

Hurricane Run 

No. 

Hurricane Rate 

(events/ year) Mean 

Standard 

Deviation Mean 

Standard 

Deviation Mean 

Standard 

Deviation 

1 1.00E-01 0.00E+00 0.00E+00 0.00E+00 0.00E+00 8.28E+07 1.66E+07 

2 5.00E-02 0.00E+00 0.00E+00 0.00E+00 0.00E+00 1.07E+08 3.31E+07 

3 1.00E-02 6.57E+07 1.22E+07 8.28E+07 1.66E+07 8.28E+07 1.66E+07 

4 1.00E-02 7.87E+07 2.11E+07 8.28E+07 1.66E+07 8.28E+07 1.66E+07 

5 1.00E-02 0.00E+00 0.00E+00 0.00E+00 0.00E+00 9.67E+07 3.30E+07 

6 1.50E-01 0.00E+00 0.00E+00 0.00E+00 0.00E+00 7.92E+07 1.89E+07 

7 5.00E-03 1.24E+08 2.39E+07 1.90E+08 3.76E+07 1.90E+08 3.76E+07 

8 9.00E-02 8.69E+07 2.99E+07 7.99E+07 1.99E+07 6.78E+06 1.79E+07 

9 1.00E-02 0.00E+00 0.00E+00 0.00E+00 0.00E+00 1.03E+08 2.43E+07 

10 5.00E-02 0.00E+00 0.00E+00 0.00E+00 0.00E+00 1.03E+08 3.21E+07 

Pumping, Rainfall and Total Water Volume in a Subbasin. The total volume entering a 

sub-basin (as a random variable with mean and standard deviation) was calculated for each 

branch of the event tree by summing volumes of water due to overtopping, breaching, and 

closure structures, as well as the water volume from rainfall and wave runup minus the effect of 

pumping. 

The pumping system in New Orleans was designed to remove rainfall from tropical storms 

up to about a 10-year event. The effect of pumping on sub-basin inflow water volumes was 

approximated by subtracting a portion of the 10-year rainfall (that considered degraded pump 

reliabilities and efficiencies) as a function of water level accumulated in a sub-basin. The water 

volume that could be pumped by a given pump station within a particular subbasin was estimated 

by taking the total individual pump station capacity and multiplying it by the duration of the 

intense portion of the rainfall for each storm. These volumes were then summed for all the 

stations within a sub-basin. This volume was considered to be the 100-percent pump station 

capacity and was subtracted from the rainfall of storm, up to the estimated 10-year rainfall 

volume. Volumes were also determined for 50-percent pump station capacity and no pump 

station capacity. 

Water Interflow between Basins and Sub-basins. Within a basin, water entering a sub-basin 

may, under certain conditions, overflow into adjacent sub-basins. Thus, prior to calculating the 

final volume of water in the sub-basins for each of the 16 branches in the event tree of Figure 9-

8, interflow among sub-basins was considered. This was done by modeling the elevations of the 

interfaces between sub-basins and determining the volume of water that would pass between 
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sub-basins for the amount of time the interface elevation was exceeded. Table 9-4 shows a 

tabulated structure for computing volumes associated with sub-basins. 

Breaching Models 

Three cases of breach failure were examined that corresponded to the breaching branches 

presented in the event tree of Figure 9- 8. The three cases are: 

1. Breach given overtopping 

2. Breach given no overtopping 

3. Breach due to feature (closure gate, pump house, etc.) or transition failures 

The first case of breach given overtopping is primarily driven by erosion, resulting from 

overtopping water flow. Fragility curves for these cases were developed as described in the 

Reliability Methodology (Appendix 10). Table 9-5 summarizes the breaching model used in the 

risk analysis. 

Breach Parameters 

The breaching scenarios require knowledge of the average breach length and depth and of the 

hydrograph at the breach location to determine basin inflows. The HPS condition after Katrina 

was reviewed to identify basic characteristics of the major breaches. The identified 

characteristics were used to develop general rules for estimating breach dimensions in the risk 

model. One critical characteristic for determining the volume of water flowing through a breach 

is the duration of time that the breach is open. During Katrina, the breaches could not be repaired 

in time to have an effect on the level of water achieved inside the basins. Therefore the time at 

which the breach occurred was assumed to have no effect on inflow volumes and water 

elevations. 

Breach without Overtopping 

IPET studies indicated that the London Ave. and 17th St. Canal breaches occurred during 

Katrina before the water level in the canals reached the top of the floodwall; the breaches 

appeared to have been the result of a foundation and/or design failure. Therefore, these breaches 

were modeled in the risk analysis as having occurred without overtopping. The high water marks 

(HWM) identified inside the Orleans Basin (where the canal breaches occurred) and the length 

of time that surge elevations exceeded lake levels in the canals were examined. The HWM 

during Katrina in the Orleans Basin was within about 1 ft of the peak surge in the canals. For 

example, it appears that the London Ave. South breach occurred when the canal water level was 

at about 7 to 8 ft, or about 3 ft below the top of wall. The peak surge in the canal was about 10 to 

11 ft, and the HWM in the Orleans Basin was about 10 ft. There was a time lag of several hours 

between the surge elevation that failed the floodwall and the peak surge elevation. This was a 

sufficient time period for the water elevation inside the Orleans Basin to reach the peak surge 
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elevation in the canal. The inverts of the canal breaches were well below the normal lake level, 

so water flowed back into the lake after the surge passed. Based on these observations, it seemed 

appropriate to use the peak surge level as the water elevation achieved inside the basin when a 

catastrophic breach (full levee height) occurred during a non-overtopping event. Therefore, for 

breaching without overtopping, the following assumptions were used in the breaching model: 

• All breaches were considered to be a result of a structural or foundation failure and the 

breach depth was set to lowest elevation of the levee or floodwall. 

• The breach depth was extended below the adjacent lake or river level. 

• The maximum basin water elevations caused by the breach were set to the maximum 

surge elevation experienced adjacent to the breach. 

Breach during an Overtopping Event 

For levees subject to overtopping and erosion, general rules were developed that determined 

breach invert elevation based on the depth of overtopping relative to the top of levee and the type 

of soil in the levee. In the case where the breach invert elevation was higher than adjacent lake or 

river levels, the depth and length of the breach, the duration of time that the surge level exceeded 

the breach invert, and the weir coefficient were required to calculate inflow water volumes for 

the breach. The breach lengths for the levees were assumed to be similar to that experienced 

during Katrina. Breach lengths at the major canal breaches varied (450 to 1000+ ft), but were all 

on the order of several hundred feet. At the industrial canal (IHNC) where overtopping did 

occur, the two Lower Ninth Ward breaches were similar in length to breaches at canals where 

overtopping did not occur. The depth of the breaches at canals where overtopping did not occur 

were below the normal canal water levels; water flowed out through these breaches when the 

surge passed. Based on these observations, it was assumed that using the peak surge level as the 

maximum water elevation achieved inside the basin was appropriate when a full-depth breach 

occurred during an overtopping event. 

For the case of a less than full-depth breach given overtopping, breach parameters for width 

and height were not available for determining inflows. The risk model did not consider breaches 

that were less than full-depth. This refinement should be added once an erosion model for levees 

subject to overtopping is available. The risk model only computed full-depth breaches. This 

approach provided a conservative estimate of basin inflows by assuming a full-depth breach. 

The following assumptions were made in the breaching events given overtopping: 

• Breaches occurred as a result of an erosion failure due to surge and/or waves. 

• All breach depths were assumed to be full levee height; however, the depth of 

overtopping required to cause a breach was dependent upon soil properties. Assumed 

values are shown in Table 9-5. 

• Durations of overtopping were calculated from the hydrographs. 
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• The maximum basin water elevations caused by the breach were set to the maximum 

surge elevation experienced adjacent to the breach. 

Table 9-5 Breaching Model 

Reaches 

Levee/Floodwall Breach Model Given Overtopping (erosion breach) 

0 to 1ft 1ft to 3ft 

Material Symbol 

Depth 

(ft) 

Breach Width (w), 

Reach Length <1000ft Depth (ft) 

Breach Width (w) (ft), 

Reach Length <1000ft Depth (ft) 

Hydraulic Fill H 0 0 9 0.50*L to max 400 18 

Clay C 0 0 3 0.50*L to max 135 13 

Unknown (Average) U 0 0 6 0.50*L to max 290 17 

Wall W 0 0 0 0 17 

Length Modifiers Reach L>1000 ft 

Overtopping Depth (ft) 

Material Symbol 0 to 1ft 1ft to 3ft >3 ft 

Hydraulic Fill H 0.0 400 < w < 0.40*L 430 < w < 0.40*L 

Clay C 0.0 135 < w < 0.10*L 135 < w < 0.10*L 

Unknown (Average) U 0.0 290 < w < 0.30*L 315 < w < 0.30*L 

Wall W 0.0 0.0 315 < w < 0.10*L 

Levee/Floodwall Breach Model Given No Overtopping 

Breach Width (w), (ft) 

Material Symbol 

Depth 

(ft) L < 1000 ft 

1000 < L< 

10,000 ft L>10,000 ft Notes 

Hydraulic Fill H 18 0.50*L to max 500 500 < w <

0.15*L 

 0.15*L 3 Breaches / 

10,000 reach 

Clay C 13 0.50*L to max 500 500 < w <

0.10*L 

0.10*L 2 Breaches / 

10,000 reach 

Unknown (Average) U 17 0.50*L to max 500 500 < w <

0.125*L 

0.125*L 2.5 Breaches / 

10,000 reach 

Wall W 17 0.50*L to max 500 500 < w <

0.075*L 

0.075*L 1.5 Breaches / 

10,000 reach 

Transitions 

Transitions Breach Model Given Overtopping 

Breach size (ft) 

Transition Type Symbol width Depth 

Ramps R 25 3 

Floodwall-Levee T 50 3 

Drainage Structures D 65 5.5 

Pump Stations P 100 5 

Gates G 25 5 

Unprotected sections U N/A N/A 

Transitions Breach Model Given No Overtopping 

Breach size (ft) 

Transition Type Symbol width Depth 

Ramps R - - Treated as opened or closed (sand bagged) 

Floodwall-Levee T - - No breaching until OT 

Drainage Structures D - - No breaching until OT 

Pump Stations P - - No breaching until OT 

Gates G - - Treat as opened or closed 

Unprotected sections U N/A N/A 
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Overtopping Volume and Rates 

The overtopping rate was computed using the rectangular weir formulae (Daugherty et al. 

1985). If the water is assumed to be an ideal liquid, it can be shown using the energy 

conservation law that the flow rate Q is given by the following equation: 

( )1/2 2 3/2 2

3 Q g LH = Eq. (9-3) 

where g is the acceleration of gravity, H is the water elevation relative to the top of the levee or 

floodwall, and L is the reach length. The actual flow rate over the weir is known to be less than 

ideal (Daugherty et al. 1985) because the effective flow area is considerably smaller than the 

product LH.

The model can be enhanced further for engineering applications by replacing the term 

( ) 2 1/2 2

3 g in Eq. 9- 3 by an empirical coefficient, known as the weir coefficient Cw, so that 

Eq. 9-3 takes on the following form: 

/ Q C LH = W3 2 Eq. (9-4) 

where 

3.33

1.84 W

if L and H are given in English units C

if L and H are given in SI units

⎧ = ⎨

⎩

Note that the Cw for the ideal fluid case is ( ) 2 1/2 2

3 g which is equal to 2.95 m/s2

. This 

coefficient is assumed to have a coefficient of variation (COV) of 0.2. Cw takes a value of 3.0, 

2.6, and 2.0 for floodwalls, levees, and gates, respectively, with a coefficient of variation of 0.2 

in English units (L and H in feet). 

For the application considered, the mean volume of the overtopping (OT) water μV for a 

given reach can be calculated as 

[ ] ( ) ∫ = C L X h t − H dt V w s s r

3 / 2

μ max ( ) ,0 Eq. (9-5) 

where a hydrograph is represented by hs(t) as illustrated in Figure 9-7; Hr is the reach height; L is 

the reach length; Cw is the weir coefficient with a coefficient of variation of 0.2, and a mean 

μ(Cw) of 3.0, 2.6, and 2.0 for floodwalls, levees, and gates, respectively; Xs is a random factor 
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with a lognormal distribution (0.20 log standard deviation and a median of 1.0). The lognormal 

distribution was applied with the following parameters: 

μ = E(ln(x)) = 0, and σ((ln(x)) = 0.2 Eq. (9-6) 

The resulting volume is the mean volume due to overtopping. The computations account for 

Xs by numerically using a step size of si Δx and n steps as follows: 

( ) ∫

∞

= −

0

3 / 2 L x h (t) H dt Vi C si s r w μ μ Eq. (9-7) 

where the probability P( si Δx ) can be computed based on the density function X s f as follows: 

s

x

P xsi f X xs dx

si

∫ s

Δ

(Δ ) = ( ) Eq. (9-8) 

such that 

( ) ( ) 1

1 1

Δ = ∑ = ∑ ∫ = = Δ

n

i

s

x

X s

n

i

P xsi f x dx

si

s Eq. (9-9) 

For each hurricane, the event tree was evaluated n times, and the branch probabilities for 

these evaluations were multiplied by the respective ) ( si P Δx according to Eq. 9-9. This step 

resulted in the number of branches produces being multiplied by n. 

The variance of the water volume for each case was computed based on the coefficient of 

variation (δ) of the weir coefficient as follows: 

2 2 ( ) σ Vi = μViδ Cw Eq. (9-10) 

where μVi is provided by Eq. 9-7, and the coefficient of variation (δ) of the weir coefficient is 

taken as 0.2. 

Failure and Overtopping Probability 

The cumulative distribution function (CDF) of the total water volume contained in a 

subbasin of n reaches was computed as follows: 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-20

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

1 i

n

V iV

i

F p F

=

=∑ Eq. (9-11) 

where pi = a overtopping probability, and FV = CDF of the total water volume. The overtopping 

probability was treated as a binary variable. For the case of point estimates of flooding per reach, 

computations were based on order statistics. Once the total volume was obtained from all 

overtopping and breach cases, the net volume (as a random variable) was computed by adding 

(or subtracting) water volumes from rainfall, wave runup and the effect of pumping. 

Event Tree Branch Probabilities 

The event tree of Figure 9-8 consists of 12 branches per hurricane. This section develops and 

summarizes the probabilities for these branches. 

The event tree includes the following primary independent sub-basin-level events: 

• C is the event that all gates within a sub-basin are closed, 

• P is the event that all pumps in the sub-basin work, and 

• B is the event that at least one reach (or one of its transition features) in a sub-basin is 

breached. 

These events were used to construct Table 9-6 that summarizes the expanded expressions for 

the probability of each branch in the event tree of Figure 9-8. Table 9-7 summarizes the 

respective procedures for water volume and elevation computation. It should be noted that the 

water volume associated with the branches involving not-all-gates closed required a procedure to 

account for all possible combinations of not-all-gates closed. Let i be the index denoting a 

unique scenario among the set of 2n

 scenarios of gate open/closed combinations (n = number of 

uncorrelated gates). The mean water volume (μ) used in the not-all-gates closed branches was: 

( ) C

i

i Ci

C

p

p

n

− =

∑=

1

2

1

μ

μ Eq. (9-12) 

where pC is the probability of all gates closed, μ Ci the mean volume associated with not-closing 

gates according to the i

th scenario, and pi the multinomial probability of the ith scenario. The 

volume variance used in the not-all-gates closed branches was: 

( )2

2

1

2 2

2

1 C

i

i Ci

C

p

p

n

− =

∑=

σ

σ Eq. (9-13) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

where 2 σ Ci is the volume variance associated with not-closing gates according to the i

th scenario. 

The subbasin interflow analysis as previously described was performed subsequent to 

Table 9-6 procedures. Water volumes were converted to elevations with a tabulated stagestorage relationship for each subbasin based on linear interpolation. Uncertainty propagation 

from the volume (V) moments ( μ V and 2 σ V ) to elevation (E) moments ( μ E and 2 σ E ) also used 

the tabulated stage-storage relationship. Linear interpolation was used since the stage-storage 

data was tabulated in increments of 1 ft. 

The results produced at this point were summarized by subbasin, for all storms and branches 

of the event tree, in the form of water elevation (mean and variance) and occurrence rate. These 

results were used to estimate an elevation-exceedance rate for a subbasin at selected elevation (e) 

values as follows: 

( ) () ( ) ( )

&

| |,

All storms branches

λ λ E >= > e P h P S h P E e hS ∑ Eq. (9-14) 

This linear relationship can be expressed as 

E a bV = + Eq. (9-15) 

where coefficients a and b were determined from interpolation. The moments of E were 

computed as 

E V μ = +a bμ Eq. (9-16) 

and 

2 22

E V σ = b σ Eq. (9-17) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-22

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 9- 6 

A Computational Summary for Branches of the Event Tree of Figure 9- 6 for a Hurricane 

and a Basin 

Branch Branch Probability (See Figure 9-6) 

1. Non-Breach 

( ) ⎛ ⎞ ∩= − ⎜ ⎟ ⎝ ⎠

()()( ) ()() 1 ( | ) () ∏ i i

i

PCPPPB O PCPP P B O P O

2. Non-Breach 

( ) ⎛ ⎞ ∩= − ⎜ ⎟ ⎝ ⎠

()()( ) ()() 1 ( | ) () ∏ i i

i

PCPPPB O PCPP P B O P O

3. Breach 

( ) ⎛ ⎞ ⎛ ⎞ ∩= − − ⎜ ⎟ ⎜ ⎟ ⎝ ⎠ ⎝ ⎠

( ) ( ) ( )1 1 ( | ) ( ) ∏ i i

i

PCPB O PC P B O P O

4. Non-Breach PCPPPB O PCPP PB PB O ()()( ) ()() () ( ) ∩ = − ∩= ( )

( ) ( ) ( ) ( ) ( ) ( ) ⎛ ⎞ ⎜ ⎟ − +− − −

⎝ ⎠

()() 1 | 1 ( | ) () 1 ( | ) () ∏ ∏ i i ii ii

i i

PCPP P B O P O P B O P O P B O P O

5. Non-Breach PCPPPB O PCPP PB PB O ()()( ) ()() () ( ) ∩ = − ∩= ( )

( ) ( ) ( ) ( ) ( ) ( ) ⎛ ⎞ ⎜ ⎟ − +− − −

⎝ ⎠

()() 1 | 1 ( | ) () 1 ( | ) () ∏ ∏ i i ii ii

i i

PCPP P B O P O P B O P O P B O P O

6. Breach PCPB O PC PB PB O ( ) ( ) ( )1 ( ) ( ) ∩= − − ∩ = ( )

( ) ( ) ( ) ( ) ( ) ( ) ⎛ ⎞ ⎜ ⎟ − − +− − −

⎝ ⎠

( )1 1 | 1 ( | ) ( ) 1 ( | ) ( ) ∏ ∏ i i ii ii

i i

PC P B O P O P B O P O P B O P O

7. Non-Breach 

( ) ⎛ ⎞ − ∩ =− − ⎜ ⎟ ⎝ ⎠

(1 ( )) ( ) ( ) (1 ( )) ( ) 1 ( | ) ( ) ∏ i i

i

PC PPPB O PC PP P B O P O

8. Non-Breach 

( ) ⎛ ⎞ − ∩ =− − ⎜ ⎟ ⎝ ⎠

(1 ( )) ( ) ( ) (1 ( )) ( ) 1 ( | ) ( ) ∏ i i

i

PC PPPB O PC PP P B O P O

m9. Breach 

( ) ⎛ ⎞ ⎛ ⎞ − ∩ =− − − ⎜ ⎟ ⎜ ⎟ ⎝ ⎠ ⎝ ⎠

(1 ( )) ( ) (1 ( )) 1 1 ( | ) ( ) ∏ i i

i

PC PB O PC P B O P O

10. Non-Breach (1 ( )) ( ) ( ) ( ) ( ) ( ) ( ) − ∩= − ∩ = PC PPPB O PCPP PB PB O ( )

( ) ( ) ( ) ( ) ( ) ( ) ⎛ ⎞ − − +− − − ⎜ ⎟ ⎝ ⎠

(1 ( )) ( ) 1 | 1 ( | ) ( ) 1 ( | ) ( ) ∏ ∏ i i ii ii

i i

PC PP P B O P O P B O P O P B O P O

11. Non-Breach (1 ( )) ( ) ( ) ( ) ( ) ( ) ( ) − ∩= − ∩ = PC PPPB O PCPP PB PB O ( )

( ) ( ) ( ) ( ) ( ) ( ) ⎛ ⎞ − − +− − − ⎜ ⎟ ⎝ ⎠

(1 ( )) ( ) 1 | 1 ( | ) ( ) 1 ( | ) ( ) ∏ ∏ i i ii ii i iPC PP P B O P O P B O P O P B O P O

12. Breach (1 ( )) ( ) ( ) 1 ( ) ( ) − ∩= − − ∩ = PC PB O PC PB PB O ( )

( ) ( ) ( ) ( ) ( ) ( ) ⎛ ⎞ − − − +− − − ⎜ ⎟ ⎝ ⎠

(1 ( )) 1 1 | 1 ( | ) ( ) 1 ( | ) ( ) ∏ ∏ i i ii ii

i i

PC P B O P O P B O P O P B O P O

Combined Branches 3 and 9 

( )( ) ( ) ⎛ ⎞ + + ∩ =− − ⎜ ⎟ ⎝ ⎠

( ) ( ) ( ) ( )( ) 1 1 ( | ) ( ) ∏ i i

i

PC PC PP PP B O P B O P O

Combined Branches 6 and 12 ( ) PC PC PP PP PB PB O () () () ()1 () ( ) + + − − ∩= ( )( )

1 1 | 1 (| ) () 1 (| ) () − − +− − − ∏ ∏ (( ( )) ( ) ( ) ) ( ) i i ii ii

i i

P B O P O PB O PO PB O PO
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-23

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 9- 7 

A Computational Summary for the Water Volumes Associated with the Branches of the 

Event Tree of Figure 9- 6 for a Hurricane and a Basin 

Branch Branch Water Volume (See Figure 9- 6) 

1. Non-Breach Use precipitation volume, and apply pumping factor 

2. Non-Breach Use precipitation volume without pumping 

3. Breach Use post-surge breach water elevation, no pumping 

4. Non-Breach Use overtopping and precipitation volume, apply pumping factor 

5. Non-Breach Use overtopping and precipitation volume without pumping 

6. Breach Use post-surge breach water elevation, no pumping 

7. Non-Breach Use precipitation and not-all-closed-closure volume apply pumping factor 

8. Non-Breach Use precipitation and not-all-closed-closure volume without pumping 

9. Breach Use post-surge breach water elevation, no pumping 

10. Non-Breach Use overtopping, precipitation and not-all-closed-closure volume apply pumping factor 

11. Non-Breach Use overtopping, precipitation and not-all-closed-closure volume without pumping 

12. Breach Use post-surge breach water elevation, no pumping 

Risk Profiles 

The construction of risk profiles required that all storms be evaluated for all possible 

combinations of events (all event tree branches) for all the basins. The number of combination 

per storm for eight basins and 12 branches of the event tree was 1,073,741,824. Dependency 

among the basins was not examined in order to reduce the number of possible combinations; 

however, the risk results obtained by examining the individual basins were considered adequate 

for evaluating the relative risks and vulnerabilities of the HPS. 

By Water Elevation

Forte results were summarized by sub-basin, for all storms and the branches of the event tree 

in the form of water elevation (mean and variance) and occurrence rate. These results were used 

to evaluate elevation-exceedance rates for a subbasin at selected elevation e values according to 

Eq. 9-18 as follows: 

> = ∑ >

All storms branches

E e P h P S h P E e h S

&

λ( ) λ ( ) ( | ) ( | , ) (9-18) 

An example of an elevation exceedence curve is shown in Figure 9-11. Given elevationexceedence probabilities and hurricane occurrence rates for a subbasin, and considering all 

storms, flood water inundation maps were developed as illustrated in Figure 9-12. The 

inundation maps show the return periods corresponding to respective elevations. 

Forte conducted interflow analyses between subbasins at the basin level for individual 

storms. The Forte analyses did not include overtopping water volumes due to wave runup, 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-24

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

pumping, or interflow analyses at the basin level when the storms were aggregated. For each 

basin, these volumes were added to the Forte water volumes using deterministic calculations. 

Three states of pumping system effectiveness were considered: no pumping, pumping at 50 

percent of capacity, and at 100 percent of capacity. The basin analyses modified the final 50, 100 

and 500 year sub-basin elevations by adding wave run-up overtopping volumes and by 

subtracting the averaged value for pumping volume expected over the entire storm set. Pumping 

volumes were estimated deterministically based on each storm's duration and intensity within 

each sub-basin, averaged over the set of storms, and subtracted from the 50, 100 and 500-yr 

elevations using the stage-storage relationship for each sub-basin. 

After modifying the basin water volumes due to wave runup and pumping, the Forte results 

at the 50, 100 and 500 year exceedence rates were examined and balanced at a basin level by 

looking at the water volumes produced at each exceedence level using the stage-storage 

relationships for the sub-basins. If the interflow elevations between subbasins were exceeded, 

water volumes were redistributed and new water surface elevations were determined. Note that 

the exceedence rates were conditional on the storm set provided to the risk team with frequencies 

as shown in Appendix 8 which do not consider tropical storms and lower intensity, more 

frequent hurricanes. The actual inundation maps developed from the final results of the risk 

analysis are shown in Appendix 13. 

By Economic and Life losses 

Using the elevation-exceedence curve, economic and life loss profiles were estimated and 

results were provided as elevation-loss curves per sub-basins. The risk profiles for the HPS are 

expressed in terms of the life loss consequences (as illustrated in Figure 9-13) and the direct 

economic (as illustrated by Figure 9-14) based upon the stage-damage curves. The stage-damage 

curves used to construct these profiles were provided by the IPET Consequence Team. The life 

and economic risk profiles developed from the final results of the risk analysis are shown in 

Appendix 13. 

Conclusions and Recommendations 

In developing the risk analysis methodology for the New Orleans hurricane protection 

system, the needs of decision and policy makers led to the requirements of producing an analytic, 

transparent, defensible, quantitative, probabilistic, and consistent methodology. Quantifying risk 

using a probabilistic framework produced elevation and loss exceedance rates based on a 

spectrum of hurricanes according the joint probability distribution of the characteristic 

parameters that define hurricane intensity and the resulting surges, waves and precipitation. The 

methodology provides a process for evaluating the performance of a hurricane protection 

systems consisting of levees, floodwalls, transitions, closure gates, drainage systems and 

pumping stations, and estimates the population and property at risk by considering the best 

estimate flood levels of each basin for occurrence rates of 0.02, 0.01, and 0.002 (i.e., average 

return periods of 50 years, 100 years, and 500 years). The quantification of risk will assist 

decision makers as they consider various alternatives to manage risk through the enhancement of 

the hurricane protection system, controlling land use, improving evacuation effectiveness, and 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-25

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

improving drainage system operations. It also provides public and private stakeholders with 

information that can be used to increase hurricane preparedness and the awareness of the risks 

associated with living in a hurricane prone environment. 

References 

ADCIRC, 2006. Finite Element Hydrodynamic Model for Coastal Oceans, Inlets, Rivers and Floodplains, 

http://www.nd.edu/~adcirc/index.htm. 

Ayyub, B. M. 2003. Risk Analysis in Engineering and Economics, Chapman & Hall/CRC Press, FL. 

Ayyub, B. M., and McCuen, R. H. 2003. Probability, Statistics and Reliability for Engineers and 

Scientists, Chapman & Hall/CRC Press, FL. 

Daugherty, R., Franzini, J., and Finnemore E., 1985, Fluid Mechanics with Engineering Applications, 598 

p., McGraw-Hill Book Co., NY. 

Dixon, T. H., Amelung, F., and Ferretti, A.. 2006. “Subsidence and Flooding in New Orleans,” Nature 

441, 587-588. 

Dokka, R. K., 2006. “Modern-Day Tectonic Subsidence in Coastal Louisiana,” Geology 34(4), 281–284. 

Eijgenraam, C. J. J., 2006. Optimal Safety Standards for Dike-Ring Areas. CPB Netherlands Bureau for 

Economic Policy Analysis: CPB Discussion Paper 62. 

Federal Emergency Management Agency (FEMA), 2006. Mitigation Assessment Team Report: Hurricane 

Katrina in the Gulf Coast - Building Performance Observations, Recommendations, and Technical 

Guidance (July 2006). http://www.fema.gov/library/viewRecord.do?id=1857 

Grossi, P. and Kunreuther, H., 2005. Catastrophe Modeling: A New Approach to Managing Risk. 

Springer-Verlag, New York. 

Holland, G. J., 1980. “An Analytic Model of the Wind and Pressure Profiles in Hurricanes.” Monthly 

Weather Review, 108, 1212-1218. 

Kumamoto, H., and Henley, E.J., 1996, Probabilistic Risk Assessment and Management for Engineers 

and Scientists, Second Edition, IEEE Press, New York. 

McGill, W. L., Ayyub, B. M., Kaminskiy, M., “Risk Analysis for Critical Asset Protection,” 

Risk Analysis, Society for Risk Analysis, (accepted for publication). 

Modarres, M., Kaminskiy, M., Krivstov, V., 1999. Reliability Engineering and Risk Analysis: A Practical 

Guide, Marcel Decker Inc., New York, NY. 

Muir-Wood, R. and Bateman, W., 2005. “Uncertainties and Constraints on Breaching and Their 

Implications for Flood Loss Estimation.” Philosophical Transactions of the Royal Society A: 

Mathematical, Physical, and Engineering Sciences, 363(1831), 1423-1430. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

National Flood Insurance Program (NFIP), 2006. Flood Insurance Manual: May 2005 (Revised October 

2006). http://www.fema.gov/business/nfip/manual200610.shtm 

US Senate, 2006. Hurricane Katrina: A Nation Still Unprepared, Report to the Committee on Homeland 

Security and Government Affairs, US Senate, Washington, DC. USACE, 2006. Interagency 

Performance Evaluation Task Force Draft Report on “Performance Evaluation of the New Orleans 

and Southeast Louisiana Hurricane Protection System,” Draft Volume VIII – Engineering and 

Operational Risk and Reliability Analysis, USACE, Washington, DC. (1 June 2006). 

https://IPET.wes.army.mil

Van Gelder, M., 2000. Statistical Methods for Risk-based Design of Civil Structures, PhD Thesis, Delft 

University of Technology, Report No. 00-1, The Netherlands. 

Van Manen, S. E., Brinkhuis, M. 2005. “Quantitative Flood Risk Assessment for Polders.” Reliability 

Engineering & System Safety, 90, 229-237. 

Voortman, H., 2003. Risk-based Design of Large Scale Flood Defence Systems, PhD Thesis, Delft 

University of Technology, Report No. 02-3, The Netherlands. 

White House, 2006. The Federal Response to Hurricane Katrina: Lessons Learned.

http://www.whitehouse.gov/reports/katrina-lessons-learned/
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 9-1. Risk Analysis Logic Diagram 

Out of Study Scope

Out of 

Paper 

Scope

Start:

Define Project 

Objectives

Hazard 

Identification & 

Analysis

Definition of 

Hurricane 

Protection System

Systems Analysis 

and Modeling

Failure Modes 

and Fragilities

Hurricane Effects 

and Rates

Vulnerability 

Analysis & 

Assessment

Water Elevation Water Elevation

Water Volume 

Computations, 

Stage-Storage, 

and Interflow

Event Tree

Elevation 

Exceedance

Water Elevation

Consequence 

Analysis: 

Economic and 

Life Loss

Economic Loss Life Loss

Uncertainty 

Analysis

Water Elevation Economic Loss Life Loss

With bounds With bounds With bounds

Decision Analysis: 

Benefit/Cost & 

Tradeoffs

Risk-Informed 

Decisions

Inundation Maps
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 9-2. Influence Diagrams for Risk Analysis 

Flooding 

(x,y)

Impacts/

Exposure

Levee/

Floodwall 

System

Mitigation Social/

Health/

Safety

Property

Environment

Culture

Subsidence 

(x,y,t)

Response/

Evacuation

Land Use

Issues

Hurricane (x,y)

(Surge, Wind,

Waves, Rain)

Closure 

Structures & 

Operations

Power

Warning

Decision/

Public 

Execution

Pump

System

Rainfall/

Tropical

Storm

Communication

Levees

Floodwall
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 9- 3. Probability Tree for the Hurricane Protection System 

Figure 9-4. Map of New Orleans and the South East Louisiana Area Showing the Geographic Bounds of 

the Study Region Considered in the Risk Analysis 

 Failure 

 Culture+Property+Environmen

 Survival 

 Culture+Property+Environmen

 Failure Mode1 

 Failure Mode2 

 Survival 

Pump

System

 Failure Mode1 

 Failure Mode2 

 Survival 

Floodwall

 Yes 

 No 

Levees

 Yes 

 No 

Communication

 Yes 

 No 

Power

 Yes 

 No 

Rainfall/

Tropical

 Failure Storm

 Success 

Warning

Decision/

Public 

Execution

 Level1 

 Level2 

 Level3 

 Others 

Closure 

Structures & 

Operations

 Mitigation 

Hurricane (x,y)

(Surge, Wind,

Waves, Rain)

Analysis of levees, 

floodwalls, etc. as a system 

 Failure 

 Culture+Property+Environmen

 Survival 

 Culture+Property+Environmen

 Failure Mode1 

 Failure Mode2 

 Survival 

Pump

System

 Failure Mode1 

 Failure Mode2 

 Survival 

Floodwall

 Yes 

 No 

Levees

 Yes 

 No 

Communication

 Yes 

 No 

Power

 Yes 

 No 

Rainfall/

Tropical

 Failure Storm

 Success 

Warning

Decision/

Public 

Execution

 Level1 

 Level2 

 Level3 

 Others 

Closure 

Structures & 

Operations

 Mitigation 

Hurricane (x,y)

(Surge, Wind,

Waves, Rain)

Analysis of levees, 

floodwalls, etc. as a system 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-30

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

1E-12

1E-10

1E-08

1E-06

0.0001

0.01

1

6 8 10 12 14 16 18

Elevation (ft)

Failure Probability

Design Elevation

Levee Top Elevation

0.5-ft Overtopping

1.0-ft Overtopping

2.0-ft Overtopping

3.0-ft Overtopping

6.0-ft Overtopping

Figure 9- 5. Hurricane Protection System Defined by Basins and Reaches 

Figure 9-6 Example of a Fragility Curve 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-31

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

#

#

Figure 9-7. Hydrographs of Storm Surge at Defined Stations in the Hurricane Protection System for a 

Single Storm Event. 

Figure 9-8. Event Tree for Quantifying Risk. 

-1

0

1

2

3

4

5

6

7

8

1 1.1 1.2 1.3 1.4 1.5 1.6 1.7

Time (days)

Elevation (ft)
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-32

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 9-9. Definition of Sub basins for New Orleans HPS. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-33

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

0.0E+00

2.0E+09

4.0E+09

6.0E+09

8.0E+09

1.0E+10

1.2E+10

1.4E+10

1.6E+10

1.8E+10

-10 -5 0 5 10 15 20

Elevation (ft)

Storage (ft^3)

Figure 9-10. Typical Stage-Storage Curve 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 9-3 

Storm Frequencies and Parameters 

Sequential 

number - 

IPET R&R 

Storm 

Storm 

Frequency 

(Events/yr) 

Central 

pressure 

deficit at 

landfall (P0) 

Radius to 

maximum 

winds at 

landfall (Rp) 

Forward 

speed at 

landfall 

(Vf) (Mph) 

Holland's 

parameter 

(B) 

Track 

angle at 

landfall 

wrt 

vertical 

(A) 

Track 

Identifer Lat Long

1 7.90E-04 960 11 11 1.27 0 1 24.43 -79.1

2 9.19E-04 960 21 11 1.27 0 1 24.43 -79.1

3 4.92E-04 960 35.6 11 1.27 0 1 24.43 -79.1

4 2.50E-03 930 8 11 1.27 0 1 24.43 -79.1

5 2.73E-03 930 17.7 11 1.27 0 1 24.43 -79.1

6 2.30E-03 930 25.8 11 1.27 0 1 24.43 -79.1

7 1.13E-03 900 6 11 1.27 0 1 24.43 -79.1

8 1.39E-03 900 14.9 11 1.27 0 1 24.43 -79.1

9 3.46E-04 900 21.8 11 1.27 0 1 24.43 -79.1

10 7.90E-04 960 11 11 1.27 0 2 24.42 -78.6

11 9.19E-04 960 21 11 1.27 0 2 24.42 -78.6

12 4.92E-04 960 35.6 11 1.27 0 2 24.42 -78.6

13 2.50E-03 930 8 11 1.27 0 2 24.42 -78.6

14 2.73E-03 930 17.7 11 1.27 0 2 24.42 -78.6

15 2.30E-03 930 25.8 11 1.27 0 2 24.42 -78.6

16 1.13E-03 900 6 11 1.27 0 2 24.42 -78.6

17 1.39E-03 900 14.9 11 1.27 0 2 24.42 -78.6

18 3.46E-04 900 21.8 11 1.27 0 2 24.42 -78.6

19 7.90E-04 960 11 11 1.27 0 3 24.42 -78.5

20 9.19E-04 960 21 11 1.27 0 3 24.42 -78.5

21 4.92E-04 960 35.6 11 1.27 0 3 24.42 -78.5

22 2.50E-03 930 8 11 1.27 0 3 24.42 -78.5

23 2.73E-03 930 17.7 11 1.27 0 3 24.42 -78.5

24 2.30E-03 930 25.8 11 1.27 0 3 24.42 -78.5

25 1.13E-03 900 6 11 1.27 0 3 24.42 -78.5

26 1.39E-03 900 14.9 11 1.27 0 3 24.42 -78.5

27 3.46E-04 900 21.8 11 1.27 0 3 24.42 -78.5

28 7.90E-04 960 11 11 1.27 0 4 24.4 -77.9

29 9.19E-04 960 21 11 1.27 0 4 24.4 -77.9

30 4.92E-04 960 35.6 11 1.27 0 4 24.4 -77.9

31 2.50E-03 930 8 11 1.27 0 4 24.4 -77.9

32 2.73E-03 930 17.7 11 1.27 0 4 24.4 -77.9

33 2.30E-03 930 25.8 11 1.27 0 4 24.4 -77.9

34 1.13E-03 900 6 11 1.27 0 4 24.4 -77.9

35 1.39E-03 900 14.9 11 1.27 0 4 24.4 -77.9

36 3.46E-04 900 21.8 11 1.27 0 4 24.4 -77.9

37 7.90E-04 960 11 11 1.27 0 5 24.43 -78.9

38 9.19E-04 960 21 11 1.27 0 5 24.43 -78.9

39 4.92E-04 960 35.6 11 1.27 0 5 24.43 -78.9

40 2.50E-03 930 8 11 1.27 0 5 24.43 -78.9

41 2.73E-03 930 17.7 11 1.27 0 5 24.43 -78.9

42 2.30E-03 930 25.8 11 1.27 0 5 24.43 -78.9
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-35

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43 1.13E-03 900 6 11 1.27 0 5 24.43 -78.9

44 1.39E-03 900 14.9 11 1.27 0 5 24.43 -78.9

45 3.46E-04 900 21.8 11 1.27 0 5 24.43 -78.9

46 3.50E-04 960 18.2 11 1.27 -45 1 24.54 -80.9

47 3.90E-04 960 24.6 11 1.27 -45 1 24.54 -80.9

48 7.16E-04 900 12.5 11 1.27 -45 1 24.54 -80.9

49 5.48E-04 900 18.4 11 1.27 -45 1 24.54 -80.9

50 3.50E-04 960 18.2 11 1.27 -45 2 24.83 -80.8

51 3.90E-04 960 24.6 11 1.27 -45 2 24.83 -80.8

52 7.16E-04 900 12.5 11 1.27 -45 2 24.83 -80.8

53 5.48E-04 900 18.4 11 1.27 -45 2 24.83 -80.8

54 3.50E-04 960 18.2 11 1.27 -45 3 25.38 -80.8

55 3.90E-04 960 24.6 11 1.27 -45 3 25.38 -80.8

56 7.16E-04 900 12.5 11 1.27 -45 3 25.38 -80.8

57 5.48E-04 900 18.4 11 1.27 -45 3 25.38 -80.8

58 3.50E-04 960 18.2 11 1.27 -45 4.1 26.08 -80.8

59 3.90E-04 960 24.6 11 1.27 -45 4.1 26.08 -80.8

60 7.16E-04 900 12.5 11 1.27 -45 4.1 26.08 -80.8

61 5.48E-04 900 18.4 11 1.27 -45 4.1 26.08 -80.8

63 2.50E-04 960 24.6 11 1.27 45 1 21.28 -90 

64 3.02E-04 900 12.5 11 1.27 45 1 21.28 -90 

65 2.01E-04 900 18.4 11 1.27 45 1 21.28 -90 

66 1.54E-04 960 18.2 11 1.27 45 2 21.3 -90 

67 2.50E-04 960 24.6 11 1.27 45 2 21.3 -90 

68 3.02E-04 900 12.5 11 1.27 45 2 21.3 -90 

69 2.01E-04 900 18.4 11 1.27 45 2 21.3 -90 

70 1.54E-04 960 18.2 11 1.27 45 3 21.27 -90.1

71 2.50E-04 960 24.6 11 1.27 45 3 21.27 -90.1

72 3.02E-04 900 12.5 11 1.27 45 3 21.27 -90.1

73 2.01E-04 900 18.4 11 1.27 45 3 21.27 -90.1

74 1.54E-04 960 18.2 11 1.27 45 4 21.28 -90 

75 2.50E-04 960 24.6 11 1.27 45 4 21.28 -90 

76 3.02E-04 900 12.5 11 1.27 45 4 21.28 -90 

77 2.01E-04 900 18.4 11 1.27 45 4 21.28 -90 

Total 7.45E-02 

62 960 18.2 11 1.27 45 1 21.28 -90 

78 960 17.7 6 1.27 0 1 24.43 -78.9

79 900 17.7 6 1.27 0 1 24.43 -78.9

80 960 17.7 6 1.27 0 2 24.42 -78.4

81 900 17.7 6 1.27 0 2 24.42 -78.4

82 960 17.7 6 1.27 0 3 24.42 -78.3

83 900 17.7 6 1.27 0 3 24.42 -78.3

84 960 17.7 6 1.27 0 4 24.4 -77.7

85 900 17.7 6 1.27 0 4 24.4 -77.7

86 960 17.7 6 1.27 0 5 24.42 -78.7

87 900 17.7 6 1.27 0 5 24.42 -78.7

88 930 17.7 6 1.27 -45 1 26.94 -80.9

89 930 17.7 6 1.27 -45 2 27.09 -80.9

90 930 17.7 6 1.27 -45 3 27.52 -80.9
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-36

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

91 930 17.7 6 1.27 -45 4.1 28.21 -80.9

92 930 17.7 6 1.27 45 1 20.66 -92.3

93 930 17.7 6 1.27 45 2 20.75 -92.6

94 930 17.7 6 1.27 45 3 20.91 -92.8

95 930 17.7 6 1.27 45 4 21.17 -93 

96 930 17.7 17 1.27 0 1 24.43 -79.1

97 930 17.7 17 1.27 0 2 24.42 -78.6

98 930 17.7 17 1.27 0 3 24.42 -78.5

99 930 17.7 17 1.27 0 4 24.4 -77.8

100 930 17.7 17 1.27 0 5 24.43 -78.9

101 930 17.7 17 1.27 -45 1 23.29 -80.8

102 930 17.7 17 1.27 -45 2 23.68 -80.9

103 930 17.7 17 1.27 -45 3 24.27 -80.8

104 930 17.7 17 1.27 -45 4.1 24.94 -80.7

105 930 17.7 17 1.27 45 1 21.28 -90 

106 930 17.7 17 1.27 45 2 21.27 -90.1

107 930 17.7 17 1.27 45 3 21.27 -90.1

108 930 17.7 17 1.27 45 4 21.26 -90.1

109 960 17.7 11 1.27 0 1.5 24.42 -78.8

110 900 17.7 11 1.27 0 1.5 24.42 -78.8

111 960 17.7 11 1.27 0 2.5 24.42 -78.5

112 900 17.7 11 1.27 0 2.5 24.42 -78.5

113 960 17.7 11 1.27 0 3.5 24.41 -78.3

114 900 17.7 11 1.27 0 3.5 24.41 -78.3

115 960 17.7 11 1.27 0 4.5 24.43 -79.1

116 900 17.7 11 1.27 0 4.5 24.43 -79.1

117 960 17.7 11 1.27 -45 1.5 24.76 -81.2

118 960 17.7 11 1.27 -45 1.5 25.15 -81.1

119 960 17.7 11 1.27 -45 2.5 25.79 -81.2

120 900 17.7 11 1.27 -45 2.5 24.76 -81.2

121 900 17.7 11 1.27 -45 3.5 25.15 -81.1

122 900 17.7 11 1.27 -45 3.5 25.79 -81.2

123 960 17.7 11 1.27 45 1.5 21.29 -90 

124 900 17.7 11 1.27 45 1.5 21.29 -90 

125 960 17.7 11 1.27 45 2.5 21.29 -90 

126 900 17.7 11 1.27 45 2.5 21.29 -90 

127 960 17.7 11 1.27 45 3.5 21.28 -90 

128 900 17.7 11 1.27 45 3.5 21.28 -90 

129 960 17.7 6 1.27 0 1.5 24.42 -78.6

130 900 17.7 6 1.27 0 1.5 24.42 -78.6

131 960 17.7 6 1.27 0 2.5 24.42 -78.3

132 900 17.7 6 1.27 0 2.5 24.42 -78.3

133 960 17.7 6 1.27 0 3.5 24.41 -78.1

134 900 17.7 6 1.27 0 3.5 24.41 -78.1

135 960 17.7 6 1.27 0 4.5 24.43 -78.9

136 900 17.7 6 1.27 0 4.5 24.43 -78.9

137 930 17.7 6 1.27 -45 1.5 26.93 -81.3

138 930 17.7 6 1.27 -45 2.5 27.23 -81.2

139 930 17.7 6 1.27 -45 3.5 27.79 -81.2
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-9-37

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

140 930 17.7 6 1.27 45 1.5 20.71 -92.5

141 930 17.7 6 1.27 45 2.5 20.83 -92.7

142 930 17.7 6 1.27 45 3.5 21.04 -92.9

143 930 17.7 17 1.27 0 1.5 24.42 -78.8

144 930 17.7 17 1.27 0 2.5 24.42 -78.5

145 930 17.7 17 1.27 0 3.5 24.41 -78.2

146 930 17.7 17 1.27 0 4.5 24.43 -79.1

147 930 17.7 17 1.27 -45 1.5 23.64 -81.3

148 930 17.7 17 1.27 -45 2.5 24.08 -81.1

149 930 17.7 17 1.27 -45 3.5 23.73 -81 

150 930 17.7 17 1.27 45 1.5 21.27 -90.1

151 930 17.7 17 1.27 45 2.5 21.27 -90.1

152 930 17.7 17 1.27 45 3.5 21.27 -90.1

Table 9- 5 

A Tabulated Structure for Water Volumes for Sub basins and Basins 

Mean (ft3) StD (ft3) Mean (ft3) StD (ft3) Mean (ft3) StD (ft3) Mean (ft) StD (ft) Mean (ft3) StD (ft3)

OW1 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 1.743E+08 4.571E+06

OW2 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 4.858E+08 9.056E+06

NOE1 0.000E+00 0.000E+00 1.655E+04 3.310E+03 4.724E+02 7.162E+01 1.187E+00 5.937E-02 4.461E+08 3.157E+07

NOE2 0.000E+00 0.000E+00 3.775E+06 7.551E+05 4.977E+02 9.954E+01 1.187E+00 5.937E-02 1.109E+09 1.355E+07

NOE3 0.000E+00 0.000E+00 2.703E+06 5.406E+05 0.000E+00 0.000E+00 1.187E+00 5.937E-02 3.059E+08 5.171E+06

NOE4 0.000E+00 0.000E+00 1.550E+01 3.100E+00 5.972E+02 1.194E+02 1.187E+00 5.937E-02 8.688E+07 2.631E+06

NOE5 0.000E+00 0.000E+00 9.367E+07 1.873E+07 0.000E+00 0.000E+00 1.187E+00 5.937E-02 2.463E+09 2.281E+07

OM1 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 7.075E+08 9.807E+06

OM2 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 6.399E+08 8.787E+06

OM3 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 2.480E+08 6.962E+06

OM4 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 7.016E+07 2.248E+06

OM5 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 4.371E+08 1.257E+07

SB1 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 1.753E+08 5.671E+06

SB2 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 1.367E+06 4.737E+04

SB3 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 1.491E+08 4.839E+06

SB4 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 1.187E+00 5.937E-02 1.581E+07 2.990E+06

Overtopping Volume (V|OT) Breach Volume

Water Volume Elevation Volume Subpolder 

Number

Precipitation Closures

Rainfall Volume
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Figure 9-11 Example Elevation-Exceedence Curve 

1.00E-06

1.00E-05

1.00E-04

1.00E-03

1.00E-02

1.00E-01

1.00E+00

-1.00E+01 -5.00E+00 0.00E+00 5.00E+00 1.00E+01 1.50E+01 2.00E+01

Elevation (ft)

Exceedance Rate

OW1

OW2

OM1

JE3

JW1

JW2

JW3

JW4

PL4

PL5
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Figure 9-12. Sample Inundation Map 
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Figure 9-13 Example Life Loss –Exceedence Curve 

Figure 9-14. Example Damage-Exceedence Curve. Note: Direct Flood Damages (Millions of 2005 Dollars) 

1.00E-03

1.00E-02

1.00E-01

1.00E+00

1.00E+01

1.00E+02

1.00E+03

1.00E+04

1.00E+05

1.00E-21 1.00E-18 1.00E-15 1.00E-12 1.00E-09 1.00E-06 1.00E-03 1.00E+00

Exceedance Rate

Life Loss

5%

25%

50%

75%

95%

Mean

0

500

1,000

1,500

2,000

2,500

3,000

3,500

4,000

0.00E+00 5.00E-02 1.00E-01 1.50E-01 2.00E-01 2.50E-01

Exceedance Rate

Direct Flood Damage (2005 Dollars)

5%

Mean

95%
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Appendix 10 

Reliability Modeling 

Introduction 

Reliability analysis is that part of the risk study that leads to an evaluation of the conditional 

probability of failure (i.e., reliability) of structures, systems, and components when they are 

exposed to the loads of a hurricane. The reliability analysis had three steps: 

1. Define and characterize the structures, components, and features constituting the 

hurricane protection system (HPS) for each drainage basin. 

2. Define failure and identify failure modes and limit states for each structure, system, 

component, and feature. 

3. Assign conditional probabilities to HPS failure states for given water elevations and wave 

heights caused by hurricane conditions. 

Two conditions were analyzed for the reliability of levees, flood walls, and pumping stations: 

pre-Katrina and post-reconstruction and repair as existing in June 2007. 

Earlier appendices contain an inventory of the structures, systems, and components in each 

drainage basin that were considered in the risk analysis. 

Approach 

The reliability of the hurricane protection system under potential water elevations due to 

surge and waves was quantified using structural and geotechnical reliability models integrated 

within a larger system description of each drainage basin. The reliability models for the HPS 

components were developed based on design and construction information, and on the results of 

the Performance Team (IPET v.V 2006) and the Pump Stations Team (IPET v.VI 2006) studies. 

Reliability models were developed and evaluated to determine dominant, or most likely, failure 

modes for each reach in each drainage basin. 

The reliability models included uncertainties in structural material properties, geotechnical 

engineering properties, subsurface soil profile conditions, and engineering performance models 
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of levees, floodwalls, and features. Uncertainties due to spatial and temporal variation, and those 

due to limited knowledge are tracked separately in the analysis, providing a best estimate of the 

frequency of failure under given loads, along with a measure of the uncertainty in that frequency. 

Physical system definition 

The HPS comprises levees, floodwalls, levees with floodwalls on top, and various points of 

transition or localized features such as pumping stations, drainage works, pipes penetrating the 

HPS, and gates. Each drainage basin perimeter was divided into segments, referred to as reaches, 

which were deemed to be homogeneous in three respects: structural cross-section, elevations in 

the cross-section, and geotechnical cross-section. In total, 135 such reaches were identified 

across all of the drainage basins, and 197 point features. 

Geometric and engineering material properties were identified for each reach and 

summarized in systems definition tables (Table 10-2). Structural cross-sections were initially 

identified by review of as-built drawings, aerial photographs, and GIS overlays; and were 

subsequently confirmed in on-site reconnaissance. Elevations were assessed in the same 

reconnaissance, supplemented by LIDAR and field surveys provided to the Risk Team. 

Geotechnical cross-sections and corresponding soil engineering properties were derived from 

original USACE General Design Memoranda (GDM) for the respective project areas of each 

drainage basin, supplemented by site characterization data collected post-Katrina at levee and 

flood wall failure sites (cone penetrometer and laboratory measurements on undisturbed 

samples). GDM’s are available in PDF format at the IPET Project web site 

(https://IPET.wes.army.mil). 

Table 10-1. Components in the Hurricane Protection System. 

1 Levees a. Embankment section (Reaches were defined on the basis of physical discontinuities - geometric, physical, 

soils, and construction characteristics) 

b. Levee Foundation 

2 Walls a. Wall structure (I or T-wall) 

b. Wall foundation 

3 Point Features a. Gates 

b. Position – open or closed 

4 Transitions a. Wall-Levee and Levee-Wall 

b. Gates 

c. Pump stations 

d. Drainage Structures 

e. Ramps 

f. Unprotected reaches 

Engineering performance models and calculations were adapted from the GDM’s. 

Engineering parameter and model uncertainties were propagated through those calculations to 

obtain approximate fragility curves as a function of water height for components of the HPS. 

These results were calibrated against the analyses of the Performance Team, which applied more 

sophisticated analysis techniques to similar structural and geotechnical profiles in the vicinity of 

failures. Failure modes identified by the Performance Team were incorporated into the reliability 

analyses as those results became available. 
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Reliability assessments were performed for individual reaches of the HPS for given water 

elevations. This resulted in fragility curves for each reach by mode of failure. 

Fragility curve. A fragility curve gives the probability of failure, conditional upon 

an event (water elevation in this study), at which a limiting failure state is 

exceeded. 

Reliability assessments for each reach and component of the drainage basin perimeter were 

combined in the HPS risk model. The risk model (Appendix 9) used water elevations from the 

hurricane hazard and the HPS fragilities to calculate probability of volume and duration of 

flooding within each drainage basin. The system risk model is structured around an event-tree 

description of the occurrence of hurricane events, corresponding water and wave heights, and the 

resulting response of the HPS. The risk model separately tracks natural variations and knowledge 

uncertainties from both the hurricane hazard and the structural and geotechnical response, to give 

a best estimate of frequency and duration of flooding, along with measures of uncertainty in 

those frequencies. 

Uncertainties and probabilities of failure 

The Corps of Engineers Technical Letter ETL 1110-2-556, “Risk-Based Analysis in 

Geotechnical Engineering for Support of Planning Studies,” (USACE 1999) suggests that the 

principal sources of uncertainty in predictions of levee performance requiring evaluation are: 

1. Uncertainty in loadings. Loadings such as floods, earthquakes, and impacts are random 

events for which magnitude and time of occurrence may be modeled by probabilistic 

methods. This often involves the use of binomial or Poisson distributions fit to observed 

event data. 

2. Uncertainty in parameter values. Geotechnical parameters such as soil strength and 

permeability have several components of uncertainty. The value of a parameter at any 

point and the average value over any distance are inherently uncertain because of soil’s 

natural spatial variability. Secondly, there is uncertainty due to testing errors and 

uncertainty in estimating the mean and variance of the properties due to the finite number 

of tests performed. The normal and lognormal distributions are often used to model 

parameter values, which may be the value at a point or the spatially averaged value 

calculated over some distance or area. 

3. Uncertainty in analytical models. Analytical models such as slope stability analysis 

methods, seepage equations, etc., have an inherent model uncertainty arising from the 

fact that they are mathematical simplifications of more complex problems, and 

unsatisfactory performance such as slope instability or piping may occur in the prototype 

at factors of safety above or below the limit state FS = 1.0 corresponding to these 

conditions in the model. Model uncertainty has not been systematically considered in 

most Corps studies to date. Where probabilistic methods are used to make economic 

comparisons of alternatives, probability values calculated using consistent models should 

provide a consistent basis for comparison even though model uncertainty is not included. 
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4. Uncertainty in performance. As parameter values and analytical models both have 

inherent uncertainty, the performance of a structure with respect to some quantifiable 

performance mode (slope stability, seepage, settlement, etc.) is likewise uncertain. The 

probability of satisfactory or unsatisfactory performance for modes with well-defined 

models and parameters is often calculated using first-order, second-moment (FOSM) 

methods, such as the Taylor’s series method, which yield a reliability index or probability 

of unsatisfactory performance Pr(U). This approach quantifies uncertainty in 

performance as a function of uncertainty in parameter values and the analytical model. 

5. Performance modes without defined limit states. In some cases, engineering models 

may not be formulated to include limit states (e.g., FS = 1) and hence may not be easily 

reformulated to provide a reliability index or probability of unsatisfactory performance. 

Instead, satisfactory performance is expected to be attained by the adoption of 

experienced-based practices. An example is the design of filter materials, where 

equations can be used to design filters expected to perform adequately and prevent 

internal erosion, but there is no measure such as the factor of safety on which to base a 

mathematical procedure for comparing the relative reliability of filters. These situations 

are not directly compatible with FOSM methods. To obtain required probability values 

for these modes, one must either use frequency models based on observed events or 

judgmental values based on expert elicitation. 

6. Frequency and magnitude of physical changes or failure events. Physical conditions 

may change at some uncertain time within the lifetime modeled in a simulation. These 

may directly lead to unsatisfactory performance or may require changing the values of 

parameters in an analytical model. Examples include scour of foundations, plugging of 

well screens by incrustation, failure of well screens by corrosion, development of seepage 

windows in sheet piling, and dislodging of fill material in rock joints. The occurrence of 

such events cannot be easily predicted by a model based on physical parameters. The 

occurrence may be modeled using a frequency-based approach such as those based on the 

exponential and Weibull distributions where sufficient data exist. 

7. Condition of unseen features. The condition of unseen features is inherently uncertain. 

Examples include the effects of unknown cracks, burrows, or other defects in levees, and 

the adequacy of grout cutoffs under dams. A similar uncertain, but non-calculable, 

situation would be determining the probability of locating and plugging the source of a 

piping channel in a foundation before destructive erosion occurs. Such situations may 

contribute considerable uncertainty regarding performance but often can only be 

accounted for in a risk assessment or reliability analysis by quantifying the experience 

and judgment of experts rather than estimating uncertainty in parameters or fitting 

distributions to historical data. 

ETL 1110-2-556 goes on to observe that geotechnical problems have a number of unique 

aspects that also require consideration in reliability analyses: 

1. In geotechnical engineering, coefficients of variation are related to the variability of 

natural materials, which may need to be assessed on a site-specific basis. 
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2. Geotechnical parameters may have relatively high coefficients of variation (the value for 

the coefficient of permeability may exceed 100 percent) and may be correlated. 

3. Soil strength parameters can be defined and analyses performed in either a total stress 

context or an effective stress context. In the former, the uncertainty in strength and pore 

pressure are lumped; in the latter, they are treated separately. 

4. Soils are continuous media where properties vary from point to point, requiring 

consideration of spatial correlation. For problems such as slope stability, the location of 

the critical free body must be searched out. Furthermore, its location varies with 

parameter values, and varying parameter values (in an FOSM or Monte Carlo analysis) 

results in different free-body locations for each set of parameter values. 

5. Although one slip surface may be “critical,” a slope can fail on any of an infinite number 

of slip surfaces; hence a slope is a system of possible failure surfaces which are 

correlated to some extent. 

6. Some earth structures such as levees may be exceedingly long, such as levees which may 

be tens of miles long. These can be treated as a number of equivalent independent 

structures; however, determining the appropriate length and number is problematical, and 

the reliability of the system may be sensitive to the assumptions made. 

The reliability analysis undertaken as part of IPET attempts to incorporate all of these 

uncertainties and considerations. 

Aleatory vs. epistemic uncertainty 

In modern practice, engineering risk analysis usually incorporates uncertainties of two 

distinct types: aleatory and epistemic. 

Aleatory uncertainty is attributed to inherent randomness, natural variation, or 

chance outcomes in the physical world; in principle, this uncertainty is irreducible 

because it is assumed to be a property of nature. Aleatory uncertainty is 

sometimes called random or stochastic variability. 

Epistemic uncertainty is attributed to lack of knowledge about events and 

processes; in principle, this uncertainty is reducible because it is a function of 

information. Epistemic uncertainty is sometimes called, subjective or internal 

uncertainty, and divides into two major sub-categories: model uncertainty and 

parameter uncertainty. 

An example of the interplay of aleatory and epistemic uncertainties in practice is the flood 

frequency curve (USACE 1998). The flood frequency curve describes natural variability or 

aleatory uncertainty of flood flows, while error bands about the curve describe epistemic 

uncertainty in the parameters of the flood frequency model. The frequency curve reflects the 

irresolvable variation of nature. The error bands reflect limited knowledge about the statistically 
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estimated parameters of the frequency curve. Collecting more data would improve our estimates 

of the parameters, and thus reduce the error bands about the frequency curve, but no amount of 

data can reduce the underlying probability distribution represented by the exceedance curve. 

Separating uncertainty into aleatory and epistemic parts is a modeling decision. Consider 

drawing a flexible curve through a set of data. A high-order curve may fit the data closely, but 

the uncertainty in the parameters of the curve will be large because there are many parameters to 

be estimated. In contrast, a straight line may not fit the data as closely, but the uncertainty in the 

slope and intercept of the line will be small. The data scatter about the curve is aleatory; the 

uncertainty in the parameters of the curve is epistemic. 

This modeling decision on whether and how to separate aleatory and epistemic uncertainty 

has important implications in a levee safety analysis. A schematic example is given in 

Figure 10-1, in which the variability of some engineering property is shown as a function of 

location. This spatial variability of the property is divided into sections thought to be 

homogeneous, and averages (means) are estimated for each. When the variability is modeled this 

way, the variations about the respective means are assumed to be aleatory uncertainties, while 

that in the estimates of the means are assumed to be epistemic uncertainties. Additional 

information can reduce the error in the estimates of the means, but it will only better characterize 

the variance about the means within each zone and not reduce that variance itself. This standard 

model of spatial variation has many implications, which are discussed in greater detail in 

Hartford and Baecher (2004). 

Figure 10-1. Schematic diagram showing the variability of some engineering property in space (e.g., soil 

strength, surge elevation, etc.). The spatial variability is divided into sections assessed to be 

homogeneous, and means are estimated for each. 

Aleatory and epistemic uncertainty affect the outcomes of a reliability analysis in different 

ways. Aleatory uncertainty manifests as variations, or frequencies of occurrence, over space or 

time. Epistemic uncertainties manifest as statistical error and systematic biases in probability 

estimates, and may introduce correlations among aleatory frequencies. 

Reach definitions 

The HPS was divided into 135 reaches. A reach is defined for the purpose of the reliability 

analysis as, 
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Reach. A continuous length of levee or wall exhibiting homogeneity of 

construction, geotechnical conditions, hydrologic and hydraulic loading 

conditions, consequences of failure, and possibly other features relevant to 

performance and risk. 

Thus, reaches are homogeneous lengths of levee or wall that differ from neighboring reaches 

in at least one of the above properties, and which are considered internally homogeneous for the 

purposes of reliability modeling and risk analysis. 

Figure 10-2. The actual but unknown fragility curve for a section is a step function at the loading conditions 

that causes failure; this is approximated by an “S-shaped” probability curve reflecting what is 

known about the levee and loads. 

All two-dimensional sections within a reach are considered to be the same with respect to 

those properties relevant to risk and reliability; thus, the fragility of the levee (i.e., probability of 

failure as a function of load) is modeled as the same everywhere within an individual reach. 

In actuality, the fragility at a particular cross section within a reach is a step function at that 

deterministic loading condition that initiates failure (Figure 10-2). Presumably, there is such a 

failure loading condition, which if it occurs will cause failure; however, that loading condition 

varies along the length of the reach and is not precisely known before a failure occurs. The 

S-shaped fragility curve reflects uncertainty about the unique loading condition that causes 

failure at a particular location. 
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Philosophically, the uncertainty represented by the S-shaped fragility curve is epistemic; that 

is, it reflects lack of knowledge. Gathering more information, for example, by performing a proof 

loading, can dramatically change the fragility curve. This is suggested in Figure 10-3, taken from 

McDonald (2002), which shows fragility curves estimated before and after high water—that is, a 

proof loading—is observed on an embankment, under which the structure performs successfully. 

Before the loading, the fragility curve is estimated as curve B. Then a higher water load is 

successfully resisted, and the fragility curve is updated to curve C. Nothing has changed in the 

structure; only the state of knowledge about the dam has changed. It may be that the water level 

that actually causes failure is that shown in curve A; but this is unknown with certainty until 

such a loading is experienced. 

The S-shaped fragility curve reflects uncertainty about the unique loading condition that 

causes failure at a particular location. Part of the uncertainty in the S-shaped fragility curve has 

to with systematic uncertainties, such as the average soil strength or average permeability along 

the reach, or the simplifications introduced in the performance models that apply everywhere; 

but another part has to do with spatial variability within the reach. Some cross sections may be 

more fragile than others, but without detailed site characterization one can’t know exactly where 

these sections are, or how much more fragile they are than the stronger sections. This part of the 

S-shaped fragility curve also reflects epistemic uncertainty, because if more detailed information 

were available the weak spots could be identified and perhaps treated as separate reaches. But, 

the detailed information is not available, which is why those potentially weak sections are 

included in the present reach, and thus the true fragility curve varies from one location to another 

in an uncertain way. In most risk and reliability analysis, this spatial variability is modeled as if 

it were aleatory—that is, as if it were random—and stochastic models are used to characterize it. 
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Figure 10-3. Hypothetical fragility curves for an earth dam as a function of pool elevation (McDonald 

2002). 

This separation of uncertainty in how fragility curves are modeled within a single reach 

introduced a length effect caused by the way we model uncertainty. Matheron (1989) is famously 

quoted as having said, “probability is in the model, not in the world.” The systematic 

uncertainties, which cause a bias in the modeling, affect every section within the reach in the 

same way: if the mean soil permeability is underestimated at one spot it is similarly 

underestimated everywhere. The spatial variability, on the other hand, does not affect every 

section in the same way: some spots are more fragile and some are less fragile. Therefore, the 

longer the reach, the higher the probability of encountering a particularly weak variation. Note, 

this length effect is due to incomplete knowledge: collect enough information and it goes away, 

but there is seldom that much information. 

Reach information was summarized in a systems definition file, which is a flat-file data base 

summarizing physical characteristics of each reach along with eight-point fragility curves. An 

example for the first 33 defined reaches is shown as Table 10-2. 

100 101 102 103 104 105

Water Level (m)

0.0

0.2

0.4

0.6

0.8

1.0

Conditional Probability of Failure

Maximum Water Level to

Date of Producing Curve B.

Subsequent Flood Level

Curve A

Curve B Curve C

[c:\cea1.grf]
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Allowance for sea-level rise and subsidence 

In follow on risk and reliability studies, for example those addressing the 2011 (100-year) 

system, an allowance is made for a potential of two (2) feet of combined sea-level rise and 

subsidence. In the IPET risk and reliability studies of the pre-Katrina conditions and “current” 

conditions as of June 2006, no allowance is made for future sea-level rise and subsidence. 

Engineering uncertainties in the reliability model 

Four categories of engineering uncertainty were included in the reliability analysis: 

1. Geological and geotechnical uncertainties, involving the spatial distribution of soils and 

soil properties within and beneath the HPS. 

2. Structural uncertainties, involving the performance of man-made systems such as levees, 

floodwalls, and point features such as drainage pipes; and the engineering modeling of 

that performance, including geotechnical performance modeling. 

3. Erosion uncertainties, involving the performance of levees and fills around floodwalls 

during overtopping, and at points of transition between levees and floodwall, in some 

cases leading to loss of grade or loss of structural support, and consequently to breaching. 

4. Mechanical equipment uncertainties, including gates, pumps, and other operating 

systems, and human operator factors affecting the performance of mechanical equipment. 

The reliability analysis takes water elevations and wave characteristics from the hurricane 

loading conditions as given, and calculates conditional probabilities of failure for specifically 

stated water elevations. Thus, hurricane effects, wind loads, water heights, and other factors of 

the loading conditions are not considered to be uncertainties in the reliability modeling. 

Uncertainties in water elevations from hurricane conditions are convoluted with the results of the 

reliability analysis in the systems risk model to generate marginal (i.e., unconditional) 

probabilities. 

Definition of Breach 

Breaches that lead to breach of the drainage basin perimeters were associated with four 

principal failure modes: (1) levee or levee foundation failure, (2) floodwall or floodwall 

foundation failure, (3) levee or floodwall erosion caused by overtopping and wave run-up, and 

(4) failure modes associated with point features such as transitions, junctions, and closures. The 

Performance Team found no failures in the HPS which originated in structural (as opposed to 

geotechnical) failure of the I-wall or T-wall components. All documented failures at I-wall and 

T-wall locations were geotechnical in nature, with structural damage resulting from the 

geotechnical failures. 
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Each reach within the drainage basin perimeter was analyzed and tracked separately, so that 

the number of breached reaches and their location around the drainage basin perimeter was 

known for each repetition of the HPS risk model. 

The pumping system may have had a mitigating effect on the water elevation of each 

drainage basin. If the capacity of the pumping system was exceeded by the inflow volume from a 

single breach then the number and location of the breaches may not matter and the pumping 

system can be ignored in the risk analysis. If, however, the inflow volume is within the capacity 

of the pumping system to remove, then the probability that the pumps are operating must also be 

calculated. The Risk Team relied on other IPET Teams to clarify technical issues. Technical 

input from other Teams helped the Risk Team determine the level of detail with which failure 

states need to be represented. 

Table 10-2. Reach systems definition table (partial section, schematic only); showing geometric, 

material, and design properties to the left; and fragility estimates to the right as a function of 

still-water level with respect to design elevations. 

PF fragility curve 

(breach|no overtopping) PF fragility curve (breach|overtopping) 

Reach 

Length 

(ft) 

Weighte

d 

elevation 

(ft) (1) 

(NAVD88 

2004.65) 

Fragility 

water 

elevation 

(ft) (2) 

(NAVD88 

2004.65) 

Reach 

type 

Foundation 

material 

type (H, C, 

S) 

Polder 

referenc

e 

Subpolder 

reference 

(3) 

Erosion 

Modifier 

for W/L 

Minimum 

Elevation 

for Pf=0 

Design 

Elev. 

(L) 6ft 

from 

TOW 

(W) 

Top of 

Levee/Top 

of Wall 

1/2 ft 

Overtopping 

1 ft 

Overtopping 

2 ft 

Overtopping

3 ft 

Overtopping

1 2,290 11.5 9.52 W H 2 NOE5 1 0.0000 0.0023 0.0034 0.0000 0.0569 0.7464 0.8787 

2 97 13.3 10.27 L H 2 NOE5 1 0.0000 0.0000 0.0000 0.0002 0.0134 0.0565 0.0658 

3 2,325 13.5 11.50 W H 2 NOE5 1 0.0000 0.0023 0.0035 0.0000 0.0094 0.5269 0.8825 

4 2,330 13.3 10.25 L H 2 NOE5 1 0.0000 0.0002 0.0003 0.0047 0.2760 0.7524 0.8049 

5 2,270 13.7 11.72 W H 2 NOE5 1 0.0000 0.0023 0.0034 0.0000 0.0184 0.6484 0.8764 

6 19,112 12.9 9.93 L H 2 NOE5 1 0.0000 0.0019 0.0029 0.0377 0.9293 1.0000 1.0000 

7 1,474 12.1 10.12 W H 2 NOE5 1 0.0000 0.0015 0.0022 0.0000 0.0370 0.5865 0.7427 

8 2,724 12.6 9.64 L H 2 NOE5 1 0.0000 0.0003 0.0004 0.0055 0.3145 0.8045 0.8520 

9 33032 18.6 15.64 L H 2 NOE5 1 0.0000 0.0033 0.0049 0.0642 0.9897 1.0000 1.0000 

10 133 18.6 15.64 L H 2 NOE1 1 0.0000 0.0000 0.0000 0.0003 0.0183 0.0766 0.0891 

11 27,665 15.1 12.13 L H 2 NOE1 1 0.0000 0.0028 0.0041 0.0541 0.9784 1.0000 1.0000 

12 8,942 16.7 13.72 L H 2 NOE1 1.05 0.0000 0.0009 0.0013 0.0187 0.7359 1.0000 1.0000 

13 7,190 17.7 14.65 L H 2 NOE1 1.1 0.0000 0.0007 0.0011 0.0158 0.6828 1.0000 1.0000 

14 22,257 15.5 12.50 L H 2 NOE1 1.1 0.0000 0.0022 0.0033 0.0480 0.9714 1.0000 1.0000 

15 111 17.5 15.50 W H 2 NOE1 1.05 0.0000 0.0001 0.0002 0.0000 0.0010 0.0624 1.0000 

16 382 20.7 18.70 W H 2 NOE2 1.05 0.0000 0.0004 0.0006 0.0000 0.0033 0.1988 1.0000 

17 10,210 16.8 13.80 L H 2 NOE2 1.1 0.0000 0.0010 0.0015 0.0223 0.8042 1.0000 1.0000 

18 10,757 17.9 14.92 L H 2 NOE2 1.1 0.0000 0.0011 0.0016 0.0000 0.0923 1.0000 1.0000 

19 9,318 20.8 18.75 W H 2 NOE3 1.05 0.0000 0.0093 0.0139 0.0000 0.2220 1.0000 1.0000 

20 7,905 17.2 14.19 L H 2 NOE3 1.1 0.0000 0.0008 0.0012 0.0173 0.7170 1.0000 1.0000 

21 539 16.7 14.72 W H 2 NOE3 1.05 0.0000 0.0005 0.0008 0.0000 0.0144 0.4758 1.0000 

22 5616 16.7 14.72 W H 2 NOE4 1.05 0.0000 0.0056 0.0084 0.0000 0.1404 0.9988 1.0000 

23 15,940 14.0 11.02 L H 2 NOE4 1.1 0.0000 0.0016 0.0024 0.0346 0.9216 1.0000 1.0000 

24 1,820 12.1 10.14 W H 2 NOE4 1.05 0.0000 0.0018 0.0027 0.0000 0.0077 0.4868 1.0000 

25 3,453 13.4 10.35 L H 2 NOE4 1.1 0.0000 0.0003 0.0005 0.0076 0.4239 1.0000 1.0000 

26 1,587 14.5 12.50 W H 2 NOE4 1.05 0.0000 0.0016 0.0024 0.0000 0.0067 0.4410 1.0000 

27 2,348 13.8 10.77 L H 2 NOE4 1.1 0.0000 0.0002 0.0004 0.0052 0.3127 1.0000 1.0000 

28 3,803 12.2 9.22 L H 2 NOE4 1.05 0.0000 0.0004 0.0006 0.0080 0.4323 0.9895 1.0000 

29 537 12.4 6.37 W H 2 NOE4 1.05 0.0000 0.0550 0.1129 0.0000 0.0023 0.1787 1.0000 

30 526 12.6 9.60 L H 2 NOE5 1.05 0.0000 0.0001 0.0001 0.0011 0.0753 0.4676 1.0000 
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Hurricane Protection System 

The HPS for each drainage basin has four components: (1) levees, (2) I-walls (which may be 

atop levees), (3) T-walls (which may be atop levees), and (4) transitions and closures. The 

reliability analysis examined the performance of the each of component, separately and in 

combination. 

The following structures in the HPS were not independently evaluated for their failure 

modes: (1) concrete aprons associated with some I-walls, and (2) sheetpiles with a short (3 to 

4 ft) concrete cap. Either could be addressed with failure modes developed for I-walls, but were 

not included in the present study. 

The following failure modes or contributing factors were not considered in the reliability 

analysis: 

1. Internal erosion (piping) of levees due to seepage; note, this is in contrast to high pore 

pressures in sand strata, which was considered, as in the vicinity of the London Avenue 

Canal or the northern end of the IHNC. Internal erosion may be reconsidered in later 

studies. 

2. The effects of maintenance on the HPS capacity over time. Improper maintenance or 

neglect can lead to reduced capacity of the levees; in particular, gates and other moving 

components also require maintenance. Trees, landscaping, and pools were observed on 

protected embankments after Hurricane Katrina, indicating a lack of code enforcement 

and maintenance of the levees. However, there was insufficient information to include 

maintenance considerations. 

3. Impact by a barge, floating debris, or other large object on the floodwalls or levees. 

4. Failure of 3-bulb water stops between I-wall sections. 

Component Performance 

For each component, a performance level was defined such that its occurrence corresponded 

to a failure to perform an intended function. The critical components within the HPS, as stated 

above, are the levees, I-walls, T-walls, and transitions and closures. These components can fail 

in a variety of modes. For each mode of failure a limit state was defined, which, if it were to 

occur, would result in a failure to keep water out of the drainage basin. 

Engineering models of the mechanics of component performance are limited in their ability 

to explicitly model a failure state. As a result, an analysis is usually carried out for incipient 

failure by examining the limits of stability. If this state is equaled or exceeded, the structure or 

component is expected to fail to perform as intended. Incipient failure models were usually 

similar to design calculations, and in many cases were adapted from the GDM’s. 

For the purpose of evaluating the performance of the levees and floodwalls, failure was 

defined as complete breaching, which allowed water to enter the drainage basin. This failure 
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occurred in two ways: (1) loss of levee or wall stability when the strength of the levee or wall 

and its foundation was insufficient to withstand the forces placed upon the structure for a given 

water elevation below the top of the wall or levee (no overtopping); or (2) overtopping caused 

the protected side of the levee or wall to erode substantially and result in a wall or levee breach, 

which allowed water to flow freely into the drainage basin. 

System Failure 

Depending on the performance of individual components in the HPS, various outcomes may 

result. For purpose of evaluating the performance of the HPS, the outcome of most interest is 

whether a protected area was flooded or not. 

The HPS was assumed to “fail” for the purposes of the risk calculations if flooding occurred 

in a protected area, beyond that expected from rainfall and runoff which can be handled by 

pumping. Given this definition, a failure of the HPS occurred even if the components making up 

the system did not fail, for example, if levees or walls were overtopped but not breached. Note, 

the main IPET report distinguishes between “engineering failures” and “breaching;” but the risk 

analysis does not. 

Flooding can occur as a result of chains of events occurring individually or in combination. 

Among these are: 

1. Levee or floodwall breaching. 

2. Inflow into an area due to levee or flood wall overtopping that does not result in 

breeching, and which exceeds the capacity of the pumping system. 

3. Inflow to an area that occurs as a result of rainfall. 

4. Inflow to an area that occurs when the capacity of the pump system is exceeded as a 

result of backflow through pump houses. 

Flooding that occurs as a result of rainfall or transient overtopping in most cases will not be 

as consequential and may be mitigated by the pumping system. 

Geological Profile 

The stratigraphy of the New Orleans area is Pleistocene and Holocene in age. Observed levee 

and floodwall failures during Katrina principally involved shallow Holocene aged sediments. 

Generally, sediments constituting the New Orleans area are less than 7,000 years old (Holocene). 

Formation of the present day New Orleans began with the rise in global sea level, beginning 

about 12,000 to 15,000 years before present. 

A typical profile for much of the New Orleans HPS shows a layer of fill at the top, underlain 

by organic clays (‘marsh’), in turn underlain by lacustrine (distributary) plastic clays, in turn 
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underlain by stiffer Pleistocene clays. Fig. 10-1 shows the profile under the New Orleans East 

(NOE) Lakefront Levee, which is typical of this profile. 

As sea level rise slowed, five short-lived delta complexes evolved across the Louisiana coast 

by depositing Mississippi River sediments through branching distributary channels. These 

channels transported and deposited fluvial sediments along the margin of the delta and built into 

shallow coastal water. Distributary channels from one of these, the St. Bernard delta, are 

responsible for filling the shallow Gulf waters in the greater New Orleans area (Frazier 1967). 

On top of these distributary clays grew cypress swamps which would eventually become the 

marsh formations. On top of these came fills, mostly clayey, on which the present levees and 

floodwalls were constructed. 

Figure 10-4. Typical geological profile, NOE lakefront section (USACE 1972). 

The spatial variability of this typical section has to do with variations in thickness of the 

various strata, and inter-bedding of sand or silt lenses and other local conditions. In some places, 

for example, the marsh can be thicker than average, as for example in the vicinity of the 17th 

Street Canal failure. 

Equally important to the performance of levees in Orleans East Bank (OEB) and NOE is the 

Pine Island Beach deposit, a buried, barrier island or beach dating to ca. 5,000 years before 

present (see Fig. 10-2). This feature extends northeast along the southern shore of Lake 

Pontchartrain, adjacent to and north of the Metairie and Gentilly ridges, former natural levees of 

the Mississippi. Sea level was 10 to 15 ft lower than the current level when the beach ridge 

formed. Consequently, foundation soils beneath OEB and NOE are affected by this buried sand, 
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which provides a high permeability channel for pore pressures. Under the London Avenue Canal 

and the northern end of the IHNC, the sand rises close to the present ground surface. 

Figure 10-5. Pine Island (buried) beach ridge, and locations of the canal breaches (after Saucier 1994). 

The 17th Street breach is located behind the axis of the beach ridge while the London Ave. 

Canal breaches are located on the axis of the ridge. Bayou Metairie is identified in red and 

forms the Bayou Sauvage distributary course. 

Soil Engineering Properties 

The principal uncertainty contributing to probability of failure of the levee and I-wall 

sections in the reliability analysis is soil engineering properties, specifically undrained shear 

strength, Su. Uncertainties in soil engineering properties are presumed to have two main 

components: (1) data scatter caused by actual variation of soil properties in space and by random 

measurement errors, and (2) systematic errors caused by limited numbers of measurements (i.e., 

statistical estimation error), and by measurement bias (see Figure 10-6). 

Uncertainty model 

The variance in soil properties is a composition of these four terms, 

 

Var(Su) = Var( x) + Var(e) + Var(m) + Var(b) (10-1) 

in which Var(⋅) is variance, Su is the soil property as input to the analysis (in this case, undrained 

strength), x is the soil property in situ, e is measurement error (noise), m is the spatial mean of 

the soil property (which has some error due to the statistical fluctuations of small sample sizes), 

and b is a model bias or calibration term caused by systematic errors in measuring soil 

engineering properties. 
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The NOE drainage basin is used here to describe the reliability analysis approach. Analyses 

of the other drainage basins are similar. The soil profile underlying NOE consists typically of 

clayey fill overlying ‘marsh’ (OH, CH), in turn overlying ‘distributary clays’ (CH), as shown in 

Figure 10-4. Critical sections in the GDMs and failures observed during Katrina occur in these 

uppermost strata. The engineering properties of deeper, stronger strata of the Pleistocene 

formations were not statistically characterized. 

Figure 10-6. Sources of soil property uncertainty in geotechnical reliability model. 

Measured Q-test results reported in the GDMs of NOE are shown as histograms in 

Figure 10-9. 
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Figure 10-7. Second-moment statistical properties of these data are shown in Table 10-3. Test values 

larger than 750 PCF were assumed to be local effects and removed from the statistics to the 

right in the table. These moments were used in subsequent calculations. 

Spatial variation 

The spatial pattern of soil variability is characterized by auto-covariance functions. These 

describe the covariance of soil properties as a function of separation distance. Soils whose 

properties vary erratically from spot to spot display little spatial covariance, while soils whose 

properties vary with more waviness display more spatial covariance. 

The auto-covariance function of a soil property z is defined as Cz(δ) = E[z(i),z(i + δ)], in 

which E[⋅] is expectation, z(i) is the soil property at some location i, and z(i+δ) is the property at 

another location at distance δ from the first. The autocorrelation function is found by 

normalizing the auto-covariance by the variance, Rz(δ) = E[z(i),z(i + δ)]Var−1

(z). The autocovariance distance is indexed as that separation distance at which Rz(δ) = e−1

. This is a 

representative or characteristic length of the spatial correlation. 

The auto-covariance function can only be estimated for distances at least as great as the 

minimum spacing among observations, that is the minimum boring spacing in the present case. 

The minimum boring spacings in NOE are on the order of many hundred feet, with some 
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spacings between adjacent borings as much as several thousand feet. To supplement the 

information in the GDMs, post-Katrina borings made in the vicinity of the 17th Street and 

London Avenue Canal breaches were used to estimate auto-covariance functions, and 

correspondingly the magnitude of measurement noise and the autocorrelation distance. 

Figure 10-8. Typical auto-covariance function for CH soils in 17th Street Canal area post-Katrina borings, 

undrained strength (PCF) from Q tests at uniform depth below grade. 

Statistical estimates of the auto-covariance were made using the ESRI Geostatistical 

Analyst®, an application running in ArcMap®. Results for the undrained strength (Q-tests) of 

London Avenue the Distributary Clay clays are shown in Figure 10-8. Analyses for Marsh and 

Fill show similar patterns. 

Table10-3. Statistics of undrained strength data (Q-tests), NOE General Design 

Memoranda. COV is the coefficient of variation, or standard deviation divided by the 

mean. 

All data Data less than 750PCF 

Parameter Fill Marsh D.Clay Fill Marsh D.Clay 

Mean (PCF) 452 405 238 333 392 238 

Std Dev (PCF) 297 154 124 142 132 124 

COV (data scatter) 0.66 0.38 0.52 0.43 0.34 0.52 
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Measurement noise 

Soil strength is measured destructively, therefore replicate measurements cannot be used to 

estimate the magnitude of random measurement error. However, the spatial covariance structure 

provides an indirect way to make the estimate (DeGroot 2006). Assuming that the measurement 

z of soil property x is corrupted by a zero-mean error e that is independent from one 

measurement to the another and independent of the value x, the measurement can be expressed 

as z=x+e. The auto-covariance function of z is the summation of the auto-covariance functions 

of x and of e: C(z)=C(x)+C(e). But, the auto-covariance function of e is a spike at the origin and 

zero otherwise. Thus, the difference between the intersection of the observed auto-covariance 

function of z extrapolated back to the origin, and the total variance Var(z), provides an estimate 

of the variance of the error, Var(e). 

The conclusions drawn from these auto-covariance analyses were: (1) the measurement noise 

(or fine-scale variation) in the Q-test data is roughly 1/2 to 3/4 the total variance of the data 

(suggesting the COVs in the top row of Table 10-4; (2) the representative auto-covariance 

distance in the horizontal direction is on the order of 1000 feet; (3) the representative autocovariance distance in the vertical direction is assumed to be on the order of 1/100 of the 

horizontal distance, or about 10 feet, although there are too few Q-test data in individual borings 

to statistically estimate this value. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-20

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 10-9. Histograms of undrained strength data (Q-tests), by soil type, NOE General Design 

Memoranda: (black) Fill, (gray) Marsh, (white) Distributary Clay. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-21

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 10-10. Histograms of undrained strength data (Q-tests), by parish, from USACE General Design 

Memoranda of various projects. 

Statistical error 

Statistical estimation error in the mean soil property is caused by limited numbers of data 

within a reach and is approximated from the standard error of sampling. The variance of the error 

is approximated as 

Var(m) ≈ Var( x)/ n , in which m is the mean soil property, x is the spatial 

variation component of data scatter, and n is the number of measurements (Table 10-4). 

Table 10-4. Estimates of component uncertainties to soil engineering property model. 

Component Fill Marsh D.Clay 

Spatial COV 0.20 0.17 0.25 

Number of measurements 48 21 23 

Statistical error in mean 0.06 0.07 0.11 

Model bias 0.1 0.1 0.1 

0%

5%

10%

15%

20%

25%

0.00

0.020.04

0.06

0.08

0.10

0.12

0.14

0.16

0.18

0.20

0.22

0.24

0.260.28

0.300.32

0.34

0.360.38

0.40

0.420.44

0.46

0.48

0.50

More

Q-Test UU (tsf)

Frequency of results .

STB

OEB

JEF

STC

NOE
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-22

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Geotechnical model error 

Model bias was calculated based on a comparison of the detailed modeling results of the 

Performance Team compared to the more simple general method of planes used in the GDMs. 

On average, the GDMs calculated factors of safety that were approximately 10% lower than 

more precise model analysis, varying from about 7% to about 18% (IPET v.V 2006). 

Fragility Curves 

Fragility curves summarize the probability of components reaching their respective limit 

states (i.e., failure), conditioned on levels of water elevation from hurricane conditions. For 

example, the fragility curve of Figure 10-11 schematically represents the probability of failure by 

deep-sliding instability of a levee section as a function of water height. Design basis water 

elevation indicates the probability of failure at the design water level (i.e., 3 ft from the top of the 

levee). 

Fragility curves for levees and floodwalls were calculated for two conditions: (1) global 

stability without overtopping, for which reliability was calculated at two water elevations, design 

elevation and top of levee, and a smooth curve approximated to lower water elevation at sea 

level; and (2) overtopping with subsequent erosion, for which reliability was estimated from 

empirical experience during Katrina at four water elevations of overtopping: 1⁄2 foot, 1 foot, 

2 feet, and 3 feet above the top of levee or flood wall. 

Once the fragility curves for each component failure mode were determined, they were input 

to the HPS risk model, which is based on event tree analysis. For each sequence in the event tree, 

a ‘sequence’ fragility curve is determined by evaluating the event tree logic at each successive 

water elevation level. Once each sequence of events has been evaluated, the composite or total 

fragility for system failure can be determined for each system performance state of interest (e.g., 

no flooding has occurred in any area protected by the HPS, or flooding occurred as a result of 

levee or floodwall failure, or flooding occurred as a result of overtopping) by simply summing 

the fragility curves for the sequence of events for the same state. 

Reliability assessments were performed for individual reaches of approximately 

homogeneous structural type, elevation, geotechnical conditions, and water elevations. This 

resulted in fragility curves for each reach by mode of failure. Such fragility curves represented 

the aleatory (i.e., random) uncertainties from one hurricane to another. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-23

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 10-5. Summary of Engineering Models used in Calculating Fragility Curves 

Failure mode Hazard 

Models and 

parameters Source of inputs Principal uncertainty 

Static instability Still-water surge; 

weak foundation 

soils 

Limiting equilibrium 

stability 

IPET v.IV, V; Soil test data; 

Design Memoranda; In situ 

surveys 

Soil properties; Still water levels; 

Existing elevations; 

Geotechnical model 

Under seepage Still-water surge; 

high permeability 

soils 

Flownet calculations; 

Limiting equilibrium 

stability 

IPET v.IV, V; Soil test data; 

Design Memoranda; In situ 

surveys 

Soil properties; Still water levels; 

Existing elevations; Geological 

profile geometry 

Still-water 

overtopping and 

scour 

Still-water surge; 

erodable fill 

Empirical correlations 

from post Katrina data 

IPET v.IV, V; Still water levels Soil fill 

properties Existing elevations 

Scour model 

Transition point 

feature erosion 

Still-water surge; 

erodable fill 

Empirical observations 

during Katrina 

IPET v.V; Still water levels Soil fill 

properties Existing elevations 

Scour model 

Wave run-up Wave heights and 

periods; erodable 

fill 

Empirical (Dutch) 

correlations and model 

test results 

IPET v.IV Wave height and period; Still 

water levels; Existing elevations 

Levee fragility, no overtopping 

Engineering performance models were adapted from the GDM for the respective reaches of 

levee. Engineering parameter and model uncertainties were propagated through those 

calculations using a first-order, second-moment approximation to obtain approximate fragility 

curves as a function of water height. The geotechnical models used in the GDMs were calibrated 

against the analysis work of the Performance Team which used more refined calculations. 

The reliability analysis was based on limiting equilibrium calculations of factor of safety 

against instability. For levees, the analysis was based on GDM calculations of factor of safety 

against wedge instability (Figure 10-12). The calculations were based on undrained (φ=0) failure 

conditions. Undrained strengths of soils underlying the levees and walls are based on Q-test 

(unconsolidated -undrained) results. 

Best estimate calculations 

Best estimate calculations were based on average (mean) soil properties, adjusted from 

calculations in the GDMs, which used factored average soil properties. That is, the calculation of 

factor of safety in the GDMs was not based on mean observed undrained strengths, but factored 

strengths, using a reduction factor of 1.2 to 1.3. These were corrected for the reliability analysis 

to yield a mean factor of safety. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-24

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 10-11. Schematic fragility curve 

Table 10-6. Soil property uncertainty by parish 

PARISH 

MEAN SU 

(TSF) 

POINT 

COV 

POINT COV 

less noise 

SPATIAL 

reduction 

COV 

Averaged N 

DEPTH of 

data 

BOUND for 

outliers 

STD DEV 

in mean 

OEB 0.24 0.43 0.22 0.70 0.15 71 0 to -40 0.4 0.05 

STB 0.16 0.53 0.27 0.70 0.19 64 0 to -40 0.4 0.07 

NOE 0.22 0.58 0.29 0.70 0.20 43 0 to -40 0.4 0.09 

STC 0.09 0.41 0.21 0.70 0.14 45 0 to -40 0.4 0.06 

JEF 0.16 0.62 0.31 0.70 0.22 169 0 to -40 0.4 0.05 

Uncertainties in undrained shear strength were propagated through the GDM calculations to 

estimate a coefficient of variation in the calculated factor of safety. The factor of safety was 

assumed to be normally distributed, and a fragility curve was approximated through a limited 

number (typically two) of calculation points. 

Soil property uncertainty in the form of coefficients of variation for undrained soil strengths 

underlying the levees and walls was propagated through the limiting equilibrium wedge stability 

calculations to obtain coefficients of variation on factors of safety, shown in Table 4. In most 

cases, the stability analyses were linear functions of undrained soil strength so that the 

coefficient of variation of the factor of safety was the same as the coefficient of variation of the 

input soil strengths. The mean factor of safety was taken as that calculated in the GDMs, 

adjusted for factored strengths. P(f|Elevation), Probability of Failure 

1.0 

0.5 

 0 

Authorization Basis 

0.15 Fractile Level 

0.85 Fractile Level 

Median 

Elevation (ft) 

Range of 

the System 

Reliability 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-25

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 10-12. Typical wedge stability analysis of levee section from GDM (USACE 1972). 

Table 10-7. Uncertainty analysis for example levee reach in NOE. 

Water level Design basis 3⁄4 design basis Top of levee 

Mean FS 1.3 2 1.2 

Spatial COV 0.17 0.17 0.17 

Spatial average reduction factor 0.8 0.8 0.8 

Systematic COV 0.07 0.07 0.07 

Total COV 0.15 0.15 0. 15 

Reliability Index, b 2.2 6 1.7 

Pf for specific 1000 ft reach 0.014 0 0.045 

Increase in Pf per 1000 feet reach 2% 0.0 5% 

Fragility curves summarize the conditional probability of levee or wall failure as a function 

of water elevation. Calculations were made for a three specific water elevations: typically design 

water level, some level lower than design (i.e. sea level), and at the top of the levee or wall. 

Uncertainty in realized factor of safety 

For a given water elevation, uncertainty in the realized factor of safety against sliding 

depends principally on the average soil strength, Su, across the area of the failure surface. This 

average strength varies from cross-section to cross-section because the soil properties themselves 

vary from spot to spot (Figure 10-13). The variability in the average soil strength is less than the 

variability in the point-to-point properties because, to some extent, the highs and lows of the soil 

strength balance against each other over the failure surface. The larger the failure surface relative 

to the autocorrelation of the soil properties, the greater the variance reduction from the local 

averages. VanMarcke (1977a,b) has shown that the variance of the spatial average for a unitwidth plain strain cross section decreases approximately in proportion to (L/rL), for L>rL, in 

which L is the cross-sectional length of the failure surface, and rL is an equivalent auto-
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-26

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

covariance distance of the soil properties across the failure surface weighted for the relative 

proportion of horizontal and vertical segments of the surface. For the wedge failure modes this is 

approximately the vertical auto-covariance distance. The variance across the full failure surface 

of width b along the axis of the levee is further reduced by averaging in the horizontal direction 

by an additional factor (b/rH), for b>rH , in which rH is the horizontal auto-covariance distance. 

At the same time that the variance of the average strength on the failure surface is reduced by the 

averaging process, so, too, the auto-covariance function of this averaged process stretches out 

from that of the point-to-point variation. 

For a failure length of approximately 1000 feet along the levee axis and 30 feet deep, with 

horizontal and vertical auto-covariance distances of 1000 feet and 10 feet, respectively, the 

corresponding variance reduction factors are approximately 0.75 for averaging over the crosssectional length L, and between 0.73 and 0.85 for averaging over the failure length b, assuming 

either an Exponential or squared-exponential (Gaussian) auto-covariance. The corresponding 

reduction to the COV of soil strength based on averaging over the failure plane is the root of the 

product of these two factors, or between 0.74 and 0.8. 

Figure 10-13. Point variation in undrained strength and variation among locally averaged strength. 

The Reliability Index for the specific levee reach of length b is the number of standard 

deviations separating the mean condition from the limiting state, 

βb = E[FS]−1

Var(FS) = E[FS]−1

ΩFSE[FS]

 (10-2) 

in which E[FS] is the mean factor of safety, Var(FS) is the variance, and ΩFS is the COV. 

Distance along levee

Undrained strength, Su .

Point-to-point varaiation

Variation in locally averaged strenth
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-27

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Seepage 

A number of seepage failure modes were considered for specific reaches of the HPS. These 

included: 

1. Increases in pore pressures in foundation or levee soils, leading to decreases in effective 

stress and thus reduced shear strength; 

2. Internal erosion (“piping”) caused by high seepage gradients in either foundation or levee 

soils, and either at the exit of the seepage or deep within the sold mass, leading to loss of 

material and the development of through-going channels; and 

3. “Blowouts” caused by pore pressures on the protected side of a levee exceeding existing 

overburden pressures caused by the weight of overlying soils, leading to voids or 

sinkholes and thus failure. 

Failures due to seepage pressures or erosion were observed during Katrina at the London 

Avenue Canal site and also present hazards at northern reaches of the IHNC. At London Avenue, 

“a line of sinkholes was observed at the inland side of the distressed east I-wall, and a sand boil 

at the inboard embankment toe indicate[d] that erosive seepage and piping had occurred beneath 

the levee” (IPET v.V 2006). In this area, the buried Pine Island Beach deposit, which is 

Holocene age (i.e., recent) sands, rises close to the present ground surface (Figure 10-5). 

The fragility curves for any reach in which these sands rise to within the critical failure zone 

under a levee or wall were adjusted for seepage pressure effects. These affected reaches included 

not only those in the vicinity of the Pine Island Sand, but also those suspected of crossing 

untreated buried stream channels in other sections of the HPS. The adjustment of the fragility 

curves was accomplished by estimating potential pore pressure rise in the affected reaches, and 

reducing effective strengths in the sand layers accordingly. This lowers the predicted mean factor 

of safety, and correspondingly increases the probability of failure at given still water levels. 

Length effect 

The HPS of New Orleans includes long lengths of embankment or wall extending many 

miles across ground that is poorly characterized from an engineering perspective. Levees fail at 

locations where loads are high and strengths are low. If these critical locations are identified 

ahead of time, traditional methods can be used to analyze stability and calculate factors of safety. 

In such situations, the overall length of levee is immaterial, because the weakest spots have been 

identified and dealt with. The probability that the levee fails is that of these weakest spots. 

The more common situation is that the levee system is not characterized with enough detail 

to know unambiguously where the weakest spots are. In this case, any reach of levee has some 

probability of experiencing higher than average loads or lower than average strengths, and as a 

result, of being a “weak spot.” Since this critical combination cannot be uniquely identified 

before a failure occurs, the longer the levee, the greater the chance that a critical combination 

exists somewhere, and thus the higher the probability of a failure somewhere. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-28

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

For a long levee, the chance of at least one failure is equivalent to the chance that the 

variations of the mean soil strength across the failure surface shown schematically in 

Figure 10-13 drop below that required for stability at least once along the length. VanMarcke 

(1977a,b) has shown that this can be determined by considering the first crossings of a random 

process. The approximation to the probability of at least one failure as provided by VanMarcke 

(1977) was used in the calculations. 

VanMarcke’s derivation is approximately as follows: For a given loading condition, that is, 

height of water, uncertainty in the realized factor of safety against sliding depends principally on 

the average soil strength, Su, across the area of the failure surface. This average strength varies 

from cross-section to cross-section because the soil properties themselves vary from spot to spot 

(Figure 10-13). The variability in the average is less than the variability in the point-to-point 

properties, because to some extent the highs and lows of strength balance out each other over the 

failure surface. The larger the failure surface relative to the autocorrelation of the soil properties, 

the more the variance of the local averages is reduced. 

VanMarcke has shown that the variance of the spatial average for a unit-width plain strain 

cross section decreases approximately in proportion to (L/rL), for L>rL, in which L is the crosssectional length of the failure surface, and rL is an equivalent autocovariance distance of the soil 

properties across the failure surface weighted for the relative proportion of horizontal and 

vertical segments of the surface. The variance across the full failure surface of width b along the 

axis of the levee is further reduced by averaging in the horizontal direction by an additional 

factor (b/rH), for b>rH , in which rH is the horizontal autocovariance distance. At the same time 

that the variance of the average strength on the failure surface is reduced by the averaging 

process, so, too, the autocovariance function of this averaged process stretches out from that of 

the point-to-point variation. 

The primary level of analysis of levee reliability is the two-dimensional levee section. The 

presumption is that this 2D section applies over a unit length of levee, defined approximately as 

the horizontal autocorrelation distance, and treated as a probabilistically independent 

characteristic length. As the total length of levee increases, the probability of systems failure 

rises in proportion to length, and soon displays a classic exponental saturation shape trending 

asypmtotically toward 1.0, according to the fomula, 

Pf =1− (1− p)

n (10-3) 

in which, Pf = 1− (1− p)

n is the probability of system failure, p is the 2D probability of failure, 

and n is the number of characteristic lengths within the reach. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-29

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Wave run-up1

Wave run-up for each reach was calculated by the approach summarized in TAW (2002). 

The average wave run-up and overtopping for levees was calculated according to the 

formula, 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = −

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = −

β

β

γ γ

ξ γ γ γ γ

γ ξ

α

m f

c

m

m b f v

c

b

m

H

R

gH

q with imum

H

R

gH

q

1

max : 0.2exp 2.6

1

exp 4.75

tan

0.067

0

3

0

0 0

0 3

0

 (10-4) 

in which, 

 q = overtopping rate [cft/s per ft] 

 g = gravitational acceleration [= 32.18 ft/s2] 

 Hm0 = significant wave height at toe of the structure [ft] surf similarity parameter [-] 

α = slope [-] 

 Rc = free crest height above still water line [ft] 

γ = influence factors for presence of berm (b), friction (f), wave incidence (β), vertical 

wall (v) 

The coefficients 4.75 and 2.6 are empirical mean values. The standard deviations aere 0.5 

and 0.35, respectively and Normally distributed. 

Equation (10-6) is valid for ξ0 < 5 and slopes steeper than 1:8. For values of ξ0 >7 the 

corresponding relationship is, 

( )⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

+ = − −

0 0

0.92

3

0

0.33 0.022

10 exp

γ f γ β m ξ

c

m H

R

gH

q

 (10-5) 

The coefficient -0.92 is the empirical mean value, with standard deviation 0.24 and Normally 

distributed. The overtopping rate in the range 5 < ξ0 < 7 is obtained by linear interpolation of 

Equations (10-6) and (10-7) using the logarithmic value of the overtopping rates. The surf 

similarity parameter is defined as: 

( )2

1,0

0

0

0

0

tan 2

−

= =

m

m

g T

H with s

s

α π ξ (10-6) 

 

1

 This section draws heavily on the internal USACE memorandum, van Ledden (2007), “Wave overtopping IPET,” 

dated 21 June. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-30

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in which, 

 s0 = wave steepness [-] 

 Tm-1,0 = mean period [s] 

Eq. (1) - (3) contain levee parameters such as the slope and the crest height and several 

influence factors for angle of wave incidence, friction, the presence of a berm and a vertical wall. 

The hydraulic levee designs in the New Orleans area generally have two steep sloping sections 

of 1:3 – 1:5 and a wave berm in between. The wave berm is located at the design still water line. 

A common wave berm factor for these levee designs is γb = 0.65 – 0.75. Apart from specific 

cases, a grass-covered levees without floodwalls on top and perpendicular wave attack is 

assumed. Hence, there is no reduction of the overtopping rate due to friction, wave attack and 

vertical walls. Based on this, we suggest to use in the risk model the following settings: slope α = 

1⁄4, a berm factor γb = 0.7 and γf = γβ = γv = 1 (“no effect”). 

The average wave overtopping over floodwalls is correspondingly, 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = −

m s

c

m H

R

gH

q

γ γβ

1 0.082exp 3.0

0

3

0

 (10-7) 

in which, 

 q = overtopping rate [cft/s per ft] 

 Hm0 = significant wave height at toe of the structure [ft] 

 Rc = free crest height above still water line [ft] 

γ = influence factors for wave incidence (β) and type of geometry (s) 

The coefficient 3.0 is the empirical mean value, with standard deviation 0.26, and a normal 

distribution assumed. The influence factors are: γs = 1 and γβ = 0.83 for plain impermeable 

floodwalls with perpendicular wave attack of short-crested waves. 

For both levees and floodwalls, the average wave overtopping can be computed using the 

still water level from ADCIRC and the wave information from STWAVE. The mean wave 

period Tm-1,0 is derived directly from the STWAVE results at 600 ft in front of the levees/ 

floodwalls1. The significant wave height at the toe of the structure (Hm0) is also derived from the 

STWAVE results, but is adapted because of depth-limited breaking in front the structure. The 

significant wave height based on the STWAVE results is limited to the maximum significant 

wave height according to: ( ) m toe H = γ ζ − z 0,max ; in which, γ : breaker parameter [-], ζ : still water 

level [ft], and ztoe : bottom level at toe of structure [ft]. 

The breaker parameter in Eq. (5) is set at γ = 0.4 in the design study. The bed level at the toe 

of most of the structures is assumed to be at ztoe = 0 ft. The standard deviation for the significant 

wave height is assumed to be 10% of the value based on STWAVE (or after reduction due to 

 

1

 Note that only the peak period Tp is available for the 152 storm suite of the 2007 situation. The peak period Tp can 

be converted easily into the mean period Tm-1,0 using Tm-1,0 = Tp/1.1. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-31

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depth-limited breaking according to Eq. (5)). The error in the wave period is set at 20% of the 

STWAVE result. The error is assumed to Normally distributed. Both errors are based on expert 

judgment due to lack of field data. 

I-Wall fragility, no overtopping 

The reliability analysis for I-walls was similarly based on limiting equilibrium calculations of 

factor of safety against instability. For I-walls, the analysis is based on the Performance Team’s 

mechanism of cracks developing in the soil immediately behind the wall and sheetpile, allowing 

hydrostatic pressure on the sheetpile. The equilibrium of a soil wedge to the protected side of the 

wall (Figure 10-14) was calculated for this condition. The calculations were based on undrained 

(φ=0) failure conditions. Undrained strengths of soils underlying the levees and walls were based 

on “Q-test” results. The design consideration of balancing forces and moments on the sheet pile 

to determine depth of penetration was considered immaterial to the reliability analysis of the wall 

sections. 

Based on the results of the Performance Team’s analyses, it was assumed that cracking 

initiated at 5 feet of water elevation on an I-wall. Thus, for water elevations lower than 5 feet, 

the factor of safety was that calculated in the GDMs. But at 5 feet, when a crack formed in the 

soil, the factor of safety underwent a step change to a forward (protected side) wedge failure. 

Figure 10-14. Sheet pile failure by deep wedge instability. 

If soil separations developed in front of the sheetpile or the levee, the condition resulted in 

increased hydrostatic forces on the flood side of the I-wall and the levee. If the separation was of 

sufficient depth, the hydrostatic forces on the wall may exceed the shear strength of the 

supporting soil and cause failure along wedge lines of least resistance behind the sheetpile. 

Reliability calculations were based on the probability that shear resistance of a wedge was 

exceeded by the loads on the levee and floodwall for a given hurricane. 
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Levee and I-Wall fragility, with overtopping 

Reliability calculations were based on the probability of overtopping causing erosion of the 

protected side of a levee that led to a breach. Two approaches were considered. The first 

approach considered flow velocities over the levee. The second approach considered water 

elevation, which is estimated by the storm surge modeling, as an indirect parameter of flow 

velocity. 

Based on the Performance Team’s results, the probability of erosion breaching was 

considered to be negligible for floodwall sections, and related to the presence of significant 

depth of hydraulic fill for levee sections. The fragility curve for levee sections was assumed to 

be that of the static failure analysis (above) up to the point of overtopping, and then a step 

function to Pf=1.0 for those sections with significant depths of hydraulic fill. For levee sections 

without significant hydraulic fill, the fragility curve remained flat at the top of wall fragility for 

overtopping. 

Figure 10-14 shows fragility curves developed for reaches defined in the Jefferson drainage 

basin. The fragilities values were developed as described in the preceding sections for conditions 

of no overtopping (water elevations up to the top of the wall or levee) and overtopping. 

Figure 10-15. Failure by rotation of I-wall, reducing I-wall elevation 

Scour and erosion 

The probability of overtopping of levees or floodwalls leading to scour and consequent 

failure was directly estimated based on empirically observed rates of failure during Katrina and 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-33

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documented in IPET Volume V (2006), and as shown in Table 10-8. These are consistent with 

later analyses by Briaud, et al. (2006). 

Table 10-8. Empirical frequency of overtopping scour failure of levees and walls as 

observed in Katrina, as a function of the velocity of overtopping flow (correlated to 

depth of overtopping) and soil type. 

LEVEES ≤0.5 FOOT ≤1.0 FOOT ≤2.0 FEET 3 FEET 

Hydraulic Fill 0 0 1 1 

Clay 0 0 0.25 0.5 

Protected 0 0 0 0.1 

WALLS ≤0.5 FOOT ≤1.0 FOOT ≤2.0 FEET 3 FEET 

Hydraulic Fill 0 0 0.5 1 

Clay 0 0 0.25 0.5 

Protected 0 0 0 0.1 

Transitions and Point Structures 

A number of HPS breaches were observed at transitions between HPS components. These 

breaches were typically at levee to I-wall, levee to T-wall or I-wall to T-wall transitions. Many 

of the HPS breaches were at point structures such as gates (road and railroad), pump stations, or 

around drainage control structures. These transitions indicate a weak link in the HPS due to the 

differing stiffness of the components which permit them to become areas of significant erosion 

during a hurricane event. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-34

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 10-16. Example of Transition Zone for East Bank of INHC 

Many of these transitions zones failed to use a “wrap-in” levee section to a more rigid wall 

structure. Instead, the levee sections sloped quickly away from the transition to expose the I- or 

T-wall. The steep slopes permitted a concentrated zone for the erosion of the levee that 

eventually exposed the I-wall or T-wall structure to additional loading and continued eroding. 

This dynamic process could lead to instability and collapse or damage to end sections of the 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-35

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

wall. An example of a levee transition for a gate section on the east bank of the INHC is shown 

in Figure 10-15. 

The failure modes for these transitions zones are complex and dynamic. The failure modes 

use the qualitative erosion parameters developed by the Performance Team as the basis for 

change in the stability of components at the transition zones. That is, the fragility of the 

transitions was taken to be similar to that of overtopped levee sections, and to depend on the 

combination of height of overtopping water and the presence of hydraulic fill enlargement to the 

levee section. Reliability for point structures (gates, control structures, pump stations) was taken 

as a point probability of failure for design loading. 

Failure Mode 1–Scour and erosion caused point structure (i.e. drainage pipe) instability. A 

levee breach may occur due to loss of the I- or T-walls at a point structure and scour could create 

instability and collapse of the structure, resulting in a breached area. 

Failure Mode 2–Breach occurs at the water stop between the I-wall and T-wall panel 

junction. This failure mode may be caused by differential displacement between panels and may 

develop tensile and shear forces in the water stop and panels. This may be due to levee erosion 

on the flood side or different rotation point between panels, or to lateral displacement of the 

levee from a foundation shear failure. This failure mode was not explicitly included in the risk 

calculations. 

Failure Mode 3–Breach at the levee and I-wall transition. This failure mode occurs due to 

levee erosion on the protected side, where the erosion starts at the end of the levee transition and 

progresses back toward the I-wall, until the I-wall rotates toward the protected side. This was 

treated as Failure Mode 1. 

Pumping Stations 

The adverse performance of mechanical, electrical, and human elements of the HPS, such as 

pumps, the availability of power, and the closure of gates, is treated as random point (i.e., 

aleatory) events with discrete probabilities of failure based on the statistical record during 

Katrina and on information provided by other IPET teams. 

The pumping stations are critical HPS system components because they maintain the flood 

levels on the protected side. Unfortunately, many of the pumping stations during Katrina reached 

and exceeded their pumping capacity shortly into the storm. Their reliability during Katrina was 

not exceedingly high as the stations primarily failed due to rising waters at the plants, a lack of 

external or backup power source, or were shut down due to inefficient pumping. These systems 

are designed to handle specific level of rainfall and are easily overwhelmed when the levees are 

overtopped by a hurricane event. The following failure modes were possible for the pumping 

stations: no commercial power, back up generator failed, mechanical fuel unavailable, pumps not 

functioning at time of incident, mechanical failure of components, operator unavailability, debris 

blocking intakes, or reversed or back flow through outfall pipes. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-36

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The reliability of the pumping stations was included in the risk model as point sources. The 

reliability is based on data collected by the Pumping Team, performance data maintained by 

Task Force Hope, and information from the dewatering plan for New Orleans developed by the 

New Orleans District. The fragility curves for each pumping station are limited to a specific 

elevation or volume of water within the drainage basin. These fragility curves vary for each 

pumping station and reflect the interior drainage areas and back flow potential as determined by 

the Interior Drainage Team. 

References 

Briaud J.-L., Chen H.-C., Govindasamy A.V., Storesund R. (2007), “Erosion tests on samples from the 

New Orleans levees”, proceedings of the Geo-Denver national conference, edited by Harold Olsen, 

ASCE, Reston, Virginia, USA. 

DeGroot, D. (2006). “Analyzing spatial variability of in situ soil properties,” Uncertainty In The 

Geological Environment, ASCE Congress, Madison v2, pp. 21-238. 

IPET (2006). Performance Evaluation of the New Orleans and Southeast Louisiana Hurricane Protection 

Systems, Final Report of the Interagency Performance Evaluation Task Force, Vol. V, 

“Performance—Levees and Floodwalls, Appendix 13, Levee erosion and scour from overtopping 

storm surge,” US Army Corps of Engineers, Washington, DC. 

IPET (2006). Performance Evaluation of the New Orleans and Southeast Louisiana Hurricane Protection 

Systems, Final Report of the Interagency Performance Evaluation Task Force, Vol. V, 

“Performance—Levees and Floodwalls, Appendix 21, Regional geology and history overview,” US 

Army Corps of Engineers, Washington, DC. 

IPET (2006). Performance Evaluation of the New Orleans and Southeast Louisiana Hurricane Protection 

Systems, Final Report of the Interagency Performance Evaluation Task Force, Vol. VI, “Interior 

Drainage and Pumping,” US Army Corps of Engineers, Washington, DC. 

IPET (2006). Performance Evaluation of the New Orleans and Southeast Louisiana Hurricane Protection 

Systems, Final Report of the Interagency Performance Evaluation Task Force, Vol. V, 

“Performance—Levees and Floodwalls, Appendix 21, Regional geology and history overview,” US 

Army Corps of Engineers, Washington, DC. 

Jack R. Benjamin & Associates (2006). “Initial technical framework paper—levee stability,” Delta Risk 

Management Strategy, Prepared for [the California] Department of Water Resources, September 6, 

18pp. 

Matheron, G. (1989). Estimating and Choosing, A. M. Hasofer, Translator, Springer Verlag, Berlin. 

McDonald, L. (2002). “Probabilities—discussion on considerations,” Working Paper, Dam Safety 

Interest Group. 

TAW (2002). “Technical report wave run-up and wave overtopping at dikes,” Technical Advisory 

Committee on Flood Defence, The Netherlands. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-10-37

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

USACE (1972). “New Orleans East Lakefront Levee Paris Road to South Point, Lake Pontchartrain,” 

Barrier Plan, DM 2 Supplement 5B, New Orleans District, June. 

USACE (1999). “Risk-Based Analysis in Geotechnical Engineering for Support of Planning Studies,” 

ETL 1110-2-556, US Army Corps of Engineers, CECW-EG, Washington, DC, 28 May. 

USACE (1998). “Risk-based analysis for flood damage reduction studies,” EM 1110-2-1619, US Army 

Crops of Engineers, Washington, DC. 

VanMarcke, E. H. (1977a). "Probabilistic modeling of soil profiles." J. Geotech. Engrg. Div., ASCE, 

103(11), 1227-1246. 

VanMarcke, E.H. (1977b). “Reliability of earth slopes,” Journal of the Geotechnical Engineering 

Division, ASCE, v.103(GT11): 1247-1265. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-11-1 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 11 

Uncertainty Analysis 

Introduction 

One of the objectives of the risk analysis was to quantitatively assess the uncertainties 

associated with the estimated levels of flooding for possible hurricane events, which included 

assessing the uncertainties associated with the occurrence rate of hurricane events, estimated 

surge and wave elevations, and the performance of the HPS (its reliability). 

There are two types of uncertainty: aleatory and epistemic uncertainty. Aleatory uncertainty 

is attributed to the inherent randomness of events in nature, manifesting as variability over time 

for phenomena that take place at a single location (temporal variability), or variability over space 

for phenomena that take place at different locations at a single time (spatial variability), or as 

variability over both time and space. These events are predicted in terms of their likelihood of 

occurring (e.g., the chance of heads in a coin flip). Aleatory uncertainty is, in principle, 

irreducible. Epistemic uncertainty is attributed to our lack of knowledge or information about 

events and processes, or our lack of understanding of physical laws that limits our ability to 

model the real world. For example, the ability to determine the likelihood of an event (i.e., its 

rate of occurrence) requires that certain data be available. If limited data are available, the 

estimated rate may be quite uncertain (i.e., statistical confidence intervals on parameter estimates 

would be large). A second type of knowledge uncertainty is attributed to a lack of understanding 

or knowledge about physical processes that must be modeled (e.g., the meteorological processes 

that generate hurricane events). In these instances expert evaluations are often required to assess 

the current state of knowledge and to quantitatively evaluate the level of uncertainty. 

Knowledge-based uncertainty is referred to as epistemic uncertainty. 

Aleatory and epistemic uncertainty affect the outcomes of a reliability analysis in different 

ways. Aleatory uncertainty manifests as variations, or frequencies of occurrence, over space or 

time. Epistemic uncertainties manifest as statistical error and systematic biases in probability 

estimates, and may introduce correlations among aleatory frequencies. 

The distinction between aleatory and epistemic uncertainty can often seem arbitrary. For 

example, the distinction depends on the models that are used in a particular analysis. 

Nonetheless, making a distinction between the sources of uncertainty in a logical manner helps 

insure that all uncertainties are quantified and identified. In principle, epistemic uncertainties are 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-11-2 

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reducible with the collection of additional data or the use/development of improved models. 

However, in a given project, it is typically not possible to reduce these uncertainties. 

Uncertainties identified as aleatory were included in the risk analysis. Aleatory uncertainty 

values were identified for the surge and wave setup elevations, rainfall, and weir coefficients. 

Epistemic uncertainties were evaluated separately from the risk analysis. 

Aleatory Uncertainties 

Surge and Wave Setup Elevation Uncertainty. The estimated surge and wave set up 

elevations were the result of wind, surge, and wave analyses conducted with PBL, ADCIRC, 

WAM, and a Boussinesq wave setup model for local effects near levees and floodwalls. These 

analyses were quite detailed, but could not include all factors that affected surge and wave setup 

elevations. 

Two types of uncertainty in the estimated water elevations provided by the hydrographs 

were: 

• random variations from location to location for a given storm 

• perfectly correlated variations for each storm, where all the hydrographs increased or 

decreased together 

The idealized characterization of hurricanes in the PBL model led to variability between the 

predicted and actual results (e.g., surge elevations). That is, if there were a set of ‘real’ 

hurricanes with the same characteristics (e.g., central pressure, etc.), the recorded hydrographs 

for each hurricane at a given location would not be exactly the same. The variability between 

modeled and actual events due to PBL modeling limitations was spatially systemic in that the 

estimated characteristics of a particular hurricane has a common effect (higher or lower surge 

elevations) at all locations. 

Factors that contributed to the uncertainty in the ADCIRC, WAM, and Bouusinesq analyses 

included time-varying tides, time-varying local topography (e.g., if trees were swept away, the 

local surface resistance to water flow would change), finite model resolution, and variations in 

the wind fields and the air/sea interaction that were not captured by the idealized PBL wind 

model. 

The detailed topography entered into the ADCIRC, WAM, and Boussinesq models 

calculated varying water elevations from location to location for each storm. The water 

elevations were dependent on the topography leading up to each location of interest on the HPS. 

Correlated variations in the water elevations for each storm were primarily due to tidal effects, 

local changes in topography over time during the event, and uncertainty in the wind modeling for 

each storm. Random variations in water elevations from location to location have a much smaller 

effect on basin flooding than do a uniform increase or decrease of the water elevation over the 

entire HPS. For this reason, local random variations in water elevations were ignored in Forte. 

Forte computed the effect of perfectly correlated variations in water elevation, as such variations 
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can have a significant effect on HPS reach failure, basin flooding, and the associated 

consequences of flooding. 

Forte has two input parameters for estimating the uncertainty in the surge and wave setup 

elevations (see App 17). However, since the surge and wave setup elevations were combined in 

the hydrographs, a single combined uncertainty value was input for the surge and wave setup 

standard. The coefficient of variation (COV) for the uncertainty in amplitude (both surge and 

wave effects) was estimated to be about 0.15, which was based on expert judgment by IPET 

team members knowledgeable about hurricane events and modeling and was applied to the entire 

duration of each hydrograph. 

Levee and Floodwall Fragility Uncertainty. Four categories of uncertainty were included 

in the HPS reliability analysis: 

1. Geological and geotechnical uncertainties, involving the spatial distribution of soils and 

soil properties within and beneath the HPS and geotechnical parameters. 

2. Structural uncertainties, involving the performance of man-made systems such as levees, 

floodwalls, and point features such as drainage pipes; and the engineering modeling of 

that performance, including geotechnical performance modeling. 

3. Erosion uncertainties, involving the performance of levees and fills around floodwalls 

during overtopping, and at points of transition between levees and floodwall, in some 

cases leading to loss of grade or loss of structural support, and consequently to breaching. 

4. Mechanical equipment uncertainties, including gates, pumps, and other operating 

systems, and human operator factors affecting the performance of mechanical equipment. 

The HPS reliability analysis calculated fragilities (conditional probabilities of failure) for 

defined reaches of each basin, where surge and wave setup elevation was the conditional 

parameter. Thus, uncertainties associated with hurricane effects, wind loads, water heights, and 

other factors of loading conditions were not considered in the reliability analysis. 

To determine the probability of failure of a given reach, uncertainties in hydrograph 

elevations were convoluted with the HPS fragilities to generate marginal (i.e., unconditional) 

probabilities of HPS failure that resulted in flooding of a basin. Fragility curves for levees and 

floodwalls were calculated for two conditions: 

1. global stability without overtopping, for which reliability was calculated at the design 

elevation and top of levee, and a smooth curve was approximated to the lower water 

elevation at sea level 

2. overtopping with subsequent erosion, for which reliability was estimated at four water 

elevations of overtopping: 1⁄2 foot, 1 foot, 2 feet, and 3 feet above the top of levee or flood 

wall, based on empirical data from Katrina. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-11-4 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reliability assessments were performed for individual reaches of approximately 

homogeneous structural type, elevation, geotechnical conditions, and water elevations. This 

resulted in fragility curves for each reach by mode of failure which were input to Forte (see 

Appendix 15). Such fragility curves represented the aleatory (i.e., random) uncertainties from 

one hurricane to another. 

The risk model (Appendix 9) used water elevations from the hurricane hazard and the HPS 

fragilities to calculate probability of volume and duration of flooding within each drainage basin. 

The risk model separately tracked uncertainties from both the hurricane hazard and the structural 

and geotechnical response, to give a best estimate of frequency and duration of flooding, along 

with measures of uncertainty in those frequencies. 

Forte input parameters for estimating the uncertainty in the HPS reliability included COVs 

for breach with overtopping, breach without overtopping, and closure structures (e.g. open 

gates). A fragility factor was added to Forte that allowed adjustment of the position of the 

fragility curve along the x-axis for epistemic uncertainty analysis, but this factor was not used in 

the final risk analyses. The default value of zero corresponded to no shift in the fragility curve. 

The COV for breach with and without overtopping events was based on the COV for the 

weir coefficient, which is described in the next section. Breach without overtopping events had a 

COV of 0.3 and breach with overtopping events had a COV of 0.2. The COV for Closure 

structures (gates) being left open was set to 0.2, based on empirical data from Katrina. 

Weir Coefficient Uncertainty 

For overtopping events, Forte computed the volume of water that flowed over a reach into a 

sub-basin using a weir equation and the height of the water above the top of the levee or 

floodwall. The weir equation uses an empirical coefficient, known as the weir coefficient Cw, 

3 / 2 Q = CW LH

where Q if the flow rate over a weir, L is the weir length, and H is the weir depth. 

The weir coefficient has an uncertainty to account for the highly complex and variable nature 

of flow over a levee or floodwall. Each time the weir equation was accessed in Forte for volume 

computations, a random coefficient was computed that modified the mean flow into the subbasin. 

The risk analysis subdivided the New Orleans HPS into approximately 135 reaches, based on 

similar construction features and elevations. A random realization of the weir coefficient was 

used over the entire length of a defined reach. This approach resulted in random variations in 

overtopping flow rates between different reaches, but assumed a uniform value within a given 

reach (i.e., the weir coefficient was perfectly correlated along any given reach). The COV of the 

weir coefficient was taken as 0.2, based on expert judgment by IPET Team members. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-11-5 

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Rainfall Uncertainty 

For each hurricane event, Forte computed an additional volume of water due to rainfall that 

was added to each sub-basin after the flooding from HPS overtopping or reach failures was 

computed. A simplified model of rainfall inside the basins was developed, as described in 

Appendix 8. The rainfall components evaluated were the mean rainfall contribution from the 

symmetric component of the mean rain field and the asymmetric component of the rain field. 

For hurricanes that passed to the right or near a given basin, the azimuthal dependence of the 

rainfall field was conservatively neglected. For hurricanes that passed to the left of a basin, the 

asymmetric component was accounted for by multiplying the above symmetric mean rainfall 

values by 1.5. This factor includes intensification due to land effects. 

Rainfall uncertainty in Forte was estimated with a lognormal random variable with mean 

value 1 and log standard deviation 0.697. This random factor was applied to the entire mean 

rainfall time history. In reality, rainfall intensity inside a given basin would display significant 

fluctuations in time and space, which locally could far exceed a factor of 2. However, the above 

random factor is considered adequate to reflect uncertainty of the total precipitation volume in a 

basin during the passage of a hurricane. 

Epistemic Uncertainties 

Only the aleatory uncertainties defined above were included in the Forte analyses. Epistemic 

uncertainties were considered separately from the Forte analyses. Sources of epistemic 

uncertainty are described below. 

Hurricane Frequency Uncertainty. The uncertainty in the rates of hurricane occurrence is 

primarily due to the limited historical sample size. Other sources of uncertainty include possible 

errors in the assumed form of marginal and conditional distributions (especially in the tail 

regions), and the uncertain near-future hurricane activity due to fluctuations and trends 

associated with climate changes and multi-decadal cycles. 

There is an epistemic uncertainty in the rate of hurricane occurrence associated with any 

particular hydrograph. This is primarily due to the relatively small strong-storm sample size. 

Future Climatic Effects Uncertainty. Uncertainty on the hurricane statistics in the Gulf of 

Mexico during the next 50 to100 years is dominated by multi-decadal oscillations. Specifically, 

considering that the North Atlantic is now experiencing a 50 percent higher-than normal activity 

and that this elevated activity may persist over a number of years and possibly for decades, it is 

reasonable for the next 50 to100 years to increase the average historical rate of hurricanes for the 

next 50 to100 years by 20 percent and to allow for an additional 25 percent uncertainty factor 

around this corrected rate. The latter factor includes uncertainty on the historical rate due to the 

finite observation period (16%) as well as uncertainty on the future evolution of the hurricane 

frequency (judgmentally assessed). 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-11-6 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Considering the general consensus and dissenting views on the effect of global warming on 

hurricane intensity, the historical mean pressure deficit is increased by 3 percent and an 

uncertainty factor of 5 percent is applied to the increased mean value. Since the effects of 

different factors on hurricane frequency and intensity are poorly correlated, these components of 

epistemic uncertainty were treated as independent. 

Table C1 in Appendix 8 (copied below) estimates that an upper bound for wave height and 

surge increases due to a doubling of high activity periods over a nominal 30-year cycle would be 

approximately 9 percent for a 500-year return period to 18 percent for a 25-year return period. 

Table C1. Estimated changes in extreme waves heights and surges for selected return 

periods, given a doubling of years with high hurricane activity (from App 8). 

Return Period (years) Change in Wave Height (percent) Change in surge (percent) 

25 +15 +18 

50 +13 +16 

100 +12 +15 

250 +11 +12 

500 +10 +9 

Hurricane Duration Uncertainty. There is uncertainty associated with the duration of each 

hydrograph, which is primarily related to uncertainties in the two factors that primarily control 

the sotrm duration, the storm radii and forward speed. The COV in both the storm radii and 

forward speed was estimated to be 0.15. The COV of the hydrograph duration was estimated to 

be the square root of the sum of the squares (√(0.152

 + 0.152), or about 0.2. 

The effect of duration uncertainty would be to lengthen or shorten the hydrograph and the 

duration of overtopping, and so increase or decrease the volume of inundation. Since the time for 

a breach to occur was not been included in the risk model, the effect of this factor was not 

considered. 

High Water Mark (HWM) Data Uncertainty. Care was taken when FEMA evaluated the 

quality of the high water mark data from Katrina. All HWM recorded throughout the Katrina 

flood region were rated on a scale that indicated whether the data was considered to be of high 

quality (more likely to be accurate) or a lower quality. Many of the high quality data points were 

documented with photos and/or video taken during the storm. 

Katrina Hydrograph Uncertainty. The Katrina storm hydrographs used in the risk analysis 

were provided by the IPET Storm Team. The hydrograph peak elevations were verified to be 

within 0.5 to 1.0 ft of measured HWM data at locations where HWM data was available outside 

of the basins. 

Katrina Inundation Uncertainty. High quality HWM data was selected to evaluate the 

estimated basin flood elevations. However, it was recognized that there were other contributing 

factors when comparing HWM and Forte elevations: 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-11-7 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

1. The HWMs were sample measurements based on one realization (i.e., Katrina). Sampling 

variability would be expected if another storm similar to Katrina was repeated many 

times. Such variability could be significant. 

2. The HWMs are not necessarily the maximum water levels after the water settled in an 

entire subbasin, but more likely a level which remained long enough to level stains. 

3. The HWMs were measured while the drainage and pumping systems were active. 

4. The HWMs were measured while backflow to the water bodies occurred at breaches. 

Hurricane Protection System and Inundation Uncertainty 

The HPS risk and reliability analysis evaluates the frequency of flooding as a result of 

rainfall, levee overtopping, and levee failure. As part of the risk and reliability analysis, the 

uncertainty in the analysis results will be estimated. 

• the uncertainty in the frequency that a basin will be flooded (due to any cause; 

overtopping, non-overtopping) 

• the uncertainty in the frequency of flood levels in each basin (i.e., the uncertainty in flood 

elevations for the 50, 100, and 500-year flood levels) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-1 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 12 

Consequences 

Overview 

One of the primary outputs of the risk and reliability modeling of the IPET Risk Team are 

estimates of the probability distributions of life loss and direct physical damage relating to the 

performance of the Hurricane Protection System (HPS) in the Greater New Orleans area. The 

risk was estimated for the following two scenarios: 

1. Pre-Katrina (August 28, 2005) 

2. The Current HPS (June 1, 2007). 

The Risk Team worked in close collaboration with Consequence Team to obtain estimates of 

life loss and property loss as a function of maximum inundation elevation in the 27 sub-basins 

that comprise the following 10 basins of the New Orleans HPS: 

East Bank 

1. New Orleans Metro - Orleans East Bank 

2. New Orleans East 

3. St. Bernard Parish 

4. Jefferson Parish 

5. St. Charles Parish 

6. Plaquemines 

West Bank 

1. Cataouache 

2. Westwego to Harvey Canal 
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3. Harvey Canal to Algiers Canal 

4. Algiers Canal to Hero Canal 

The numbers of sub-basins that are contained within portions of the following parishes are 

indicated in parentheses: Jefferson (7), Orleans (12), Plaquemines (1), St. Bernard (5), and St. 

Charles (2) Parishes (See Figure 12-1). 

Figure12-1: Sub-basin Map 

The Risk Model was run for different hurricane realizations that represent a wide range of 

hurricane events with different severities, directions, points of landfall, etc. For each of these 

hurricane realizations, the Risk Model represented the performance of the HPS and estimated the 

probability that inundation would result from insufficient internal drainage, overtopping of the 

levees, and levee breaching. The resulting estimates of maximum inundation depths were used as 

a basis for interpolation of life loss and property loss estimates using the relationships provided 

by the Consequence Team. Estimates were made for each of the 27 sub-basins and for the preKatrina and current scenarios. Thus, it was necessary for the life loss and property loss estimates 

to cover a range of elevations associated with a range of hurricane events that could impact New 

Orleans from minor inundation to an elevation 36 ft above sea level. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-3 

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The estimates of life loss were developed as probability distributions and the estimates of 

property loss were developed as best estimates with an associated 90 percent confidence interval 

rather than single-value or point estimates. The probability distributions for life loss and 

confidence intervals for property losses represent various types of uncertainties in the estimates, 

which are described below. 

Life Loss Estimation 

Life loss was estimated by the Consequence Team using two computer models as follows: 

• LIFESim Modeling System1

 was developed: a) to estimate how the population in the 

flooded sub-basins would redistribute vertically in relation to the depth of inundation; 

and b) to classify population into one of three flood lethality zones, which are defined in 

the LIFESim model and by McClelland and Bowles (2002), and an additional sub-zone 

for people who would be expected to be able to walk away from the inundation area 

following inundation. Thus, LIFESim was run without evacuation. 

• A Monte Carlo Uncertainty Model, which was developed to take the vertically redistributed estimates of population in the three flood lethality zones from LIFESim, 

estimated: a) the immediate loss of life using fatality rate probability distributions from 

LIFESim and McClelland (2000) accounting for evacuation effectiveness as a random 

variable that varied according to a triangular probability distribution (65 percent, 

80 percent, 95 percent); and b) delayed fatalities amongst those who survived the initial 

inundation but were not rescued, where the rescue effectiveness was accounted for a 

random variable that varied according to a uniform probability distribution between 

99.5 percent and 100 percent in the Safe Flood Lethality Zone and between 95 percent 

and 100 percent in the Compromised and Chance Flood Lethality Zones. 

The entire process is described in a report by Abt Associates, Inc. (2006) and involved the 

following steps: 

1) Calibration of LIFESim to the Hurricane Katrina event. 

2) Use of LIFESim and the Monte Carlo Uncertainty Model to estimate immediate and 

delayed life loss associated with the entire range of maximum inundation levels, given 

the population and housing stock that existed prior to Hurricane Katrina. 

3) Use of LIFESim and the Monte Carlo Uncertainty Model to estimate immediate and 

delayed life loss associated with the entire range of maximum inundation levels, given 

the population and housing stock that were expected to exist June 1, 2006. 

Various limitations in this approach and some potential future improvements are described in 

the report by Abt Associates, Inc. (2006). The distributions of life-loss estimates for each 

 

1

 Institute for Dam Safety Risk Management at Utah State University (Aboelata and Bowles 2005) LIFESim 

includes a simulation module for warning and evacuation, which was not used in this study. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-4 

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sub-basin currently combine aleatory uncertainty associated with the fatality rates in each Flood 

Lethality Zone and epistemic uncertainties associated with evacuation and rescue 

effectivenesses. It is desired to treat the epistemic and aleatory uncertainties separately in future 

work. 

Direct Economic Loss Estimation 

The objective of the direct economic damage analysis was to develop potential stage-damage 

curves that might represent the flood damage potential as of June 1, 2007. This required 

accounting for the severity of the Katrina damage and the amount of property loss recovery since 

Katrina. In some areas flooded by Katrina, where water depths were low, recovery has been 

almost complete. In other areas, where water depths were high, little recovery or reinvestment 

has taken place. It is extremely difficult and at the peril of the analyst to make general estimates 

about the amount of recovery. None the less, some guidance exists in terms of what others have 

assumed about recovery. The analysis conducted followed the basic parameters provided in the 

RAND Gulf States Policy Institute published a report titled “The Repopulation of New Orleans 

after Hurricane Katrina” (2006). In developing estimates of repopulation over time, the authors 

relied on the depth of flooding as the basic determinant of the rate of population recovery. 

Table 12-1 shows the depth categories and damage recovery rates assumed in developing the 

June 2006 stage-damage. A RAND category of <2 foot was subdivided into two categories: 

<1 foot and 1 to 2 feet. Additionally, the >4 feet category was subdivided into three categories: 4 

to 6 feet; 6 to 8 feet; and >8 feet. These categories are consistent with those used in social, 

cultural and historic analysis of the impacts of Katrina the post-Katrina recovery. However, the 

values of recovery rates are to some degree arbitrary and other rates may be justified. For the 

estimate of the post-Katrina stage-damage functions shown in this section, these rates were used. 

The use of these values resulted in an estimate of the March 2006 population of New Orleans of 

approximately 155,000 people. This is within the range of other estimates. 

The June 2006 estimate of potential stage-property damage started with these assumptions. 

However, the range of depths of flooding was expanded to include more depth of flooding 

categories while preserving the basic concept. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-5 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table12-1: Estimated Repopulation Rates for New Orleans 

Period Depth of Flooding Repopulation Rate (%) 

No flooding 65 

<2 feet 20 

2–4 feet 5 

December 2005 

>4 feet 1 

No flooding 100 

<2 feet 35 

2–4 feet 15 

March 2006 

>4 feet 5 

No flooding 110 

<2 feet 75 

2–4 feet 25 

September 2006 

>4 feet 10 

Source: RAND (2006), “The Repopulation of New Orleans after Hurricane Katrina.” 

Approach 

The post-Katrina stage-damage tables and curves were estimated using the same sub-basin 

definitions as for the pre-Katrina values. Additionally, the estimation started with the same 

census block approach. The Katrina depth grid was used to estimate the depth of flooding for 

each census block. These depths were then used to select the census blocks that incurred 

damages within each of the categories shown in Table 12-1. For instance, within the Orleans 

Metro 5 sub-basin, 1,535 census blocks had flooding of 1 foot or less while a total of 4,400 

census blocks were flooded. Table 12-2 shows the complete estimate of the number of the census 

blocks flooded by Katrina by depth category. 

From these selected census blocks, damages at each stage were aggregated to the sub-basin 

level for each of the recovery category. This calculation determined the amount of the Katrina 

damage within each depth category. This was repeated for each of the Katrina flood depth 

categories. 

For each resulting sub-basin stage-damage, the recovery factors from Table 12-1 were 

applied. The recovered potential damage value was then aggregated at each stage. This provides 

an estimate of the Current HPS potential property damage at each stage for all property damaged 

estimated to have occurred from Katrina. The last step in the process was to adjust the potential 

pre-Katrina stage-damages by first subtracting the Katrina damage at each stage and then adding 

the potential recovered damage at each stage. This was necessary because the Katrina flood 

levels were not high enough to damage all the property in a sub-basin, at least for some 

sub-basins. 

Therefore, for some property, recovery from flooding was not necessary, so it contributed its 

full damage potential to the post-Katrina, Current HPS, stage-damage. Table 12-5 provides the 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-6 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

pre-Katrina damage potential and the Current HPS damage potential by stage for all sub-basins 

flooded by Katrina. Stage-damage for all other sub-basins retained the pre-Katrina stagedamages. All stages are in NAVD88 (2004.65). 

Tables 12-3 and 12-4 provide the pre-Katrina fatality potential and the Current HPS fatality 

potential by stage for all sub-basins flooded by Katrina. Note these tables reflect the post-Katrina 

repopulation as 1 June 2006. Stage-fatality for all other sub-basins retained the pre-Katrina 

stage-fatalities. All stages are in NAVD88 (2004.65). 

Table 12-2: Number of Census Blocks within Each Subbasin Flooded by Katrina by 

Depth Category 

Count Of Census Blocks within Katrina Flood Depth Category 

Subbasin Name 0-1 feet 1 to 2 feet 2 to 4 feet 4 to 6 feet 6 to 8 feet > 8 feet 

JE2 5 6 8 1 1 1 

NOE2 1 2 2 10 19 7 

NOE3 7 8 12 8 59 7 

NOE4 18 3 0 0 0 0 

NOE5 27 31 156 173 371 99 

OM1 37 37 107 126 163 361 

OM2 24 24 46 56 121 321 

OM3 301 136 387 358 219 61 

OM4 63 51 72 50 9 1 

OM5 1535 346 871 957 640 35 

SB1 31 25 91 153 200 375 

SB3 62 32 49 117 173 44 

SB4 5 37 62 50 13 0 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-7 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality 

OW1 OW2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 8.720E-03 2.237E-01 1.010E-01 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 1.744E-02 4.473E-01 2.021E-01 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 1.188E-01 1.940E+00 8.842E-01 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 2.202E-01 3.433E+00 1.566E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 1.176E+00 6.997E+00 3.671E+00 

-3.000E+00 6.353E-05 1.630E-03 7.361E-04 2.132E+00 1.056E+01 5.776E+00 

-2.000E+00 1.271E-04 3.259E-03 1.472E-03 3.266E+00 1.810E+01 9.616E+00 

-1.000E+00 5.400E-04 1.385E-02 6.257E-03 4.399E+00 2.565E+01 1.346E+01 

0.000E+00 9.529E-04 2.444E-02 1.104E-02 6.662E+00 3.478E+01 1.871E+01 

1.000E+00 1.348E-03 3.458E-02 1.562E-02 8.925E+00 4.392E+01 2.397E+01 

2.000E+00 1.744E-03 4.472E-02 2.020E-02 1.431E+01 7.348E+01 3.785E+01 

3.000E+00 6.153E-03 1.485E-01 6.724E-02 1.969E+01 1.030E+02 5.173E+01 

4.000E+00 1.056E-02 2.523E-01 1.143E-01 4.150E+01 1.878E+02 1.009E+02 

5.000E+00 3.526E-02 7.662E-01 3.486E-01 6.332E+01 2.726E+02 1.500E+02 

6.000E+00 5.996E-02 1.280E+00 5.829E-01 1.080E+02 4.066E+02 2.434E+02 

7.000E+00 8.664E-02 1.399E+00 6.363E-01 1.526E+02 5.406E+02 3.367E+02 

8.000E+00 1.133E-01 1.519E+00 6.896E-01 1.909E+02 6.699E+02 4.200E+02 

9.000E+00 2.877E-01 2.529E+00 1.137E+00 2.292E+02 7.992E+02 5.032E+02 

1.000E+01 4.620E-01 3.540E+00 1.584E+00 2.819E+02 9.969E+02 6.231E+02 

1.100E+01 1.359E+00 6.906E+00 3.515E+00 3.347E+02 1.195E+03 7.429E+02 

1.200E+01 2.256E+00 1.027E+01 5.447E+00 4.499E+02 1.670E+03 1.015E+03 

1.300E+01 3.810E+00 1.438E+01 8.587E+00 5.652E+02 2.145E+03 1.287E+03 

1.400E+01 5.364E+00 1.849E+01 1.173E+01 8.275E+02 2.997E+03 1.850E+03 

1.500E+01 5.855E+00 2.018E+01 1.280E+01 1.090E+03 3.849E+03 2.413E+03 

1.600E+01 6.346E+00 2.188E+01 1.388E+01 1.337E+03 4.678E+03 2.948E+03 

1.700E+01 7.170E+00 2.823E+01 1.640E+01 1.584E+03 5.507E+03 3.483E+03 

1.800E+01 7.993E+00 3.457E+01 1.893E+01 1.806E+03 6.274E+03 3.970E+03 

1.900E+01 1.336E+01 6.909E+01 3.400E+01 2.027E+03 7.041E+03 4.458E+03 

2.000E+01 1.873E+01 1.036E+02 4.908E+01 2.272E+03 7.895E+03 4.998E+03 

2.100E+01 3.813E+01 1.518E+02 8.787E+01 2.516E+03 8.748E+03 5.538E+03 

2.200E+01 5.753E+01 2.000E+02 1.267E+02 2.779E+03 9.660E+03 6.118E+03 

2.300E+01 6.151E+01 2.139E+02 1.356E+02 3.042E+03 1.057E+04 6.698E+03 

2.400E+01 6.550E+01 2.278E+02 1.445E+02 3.276E+03 1.138E+04 7.214E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-8 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

NOE1 NOE2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 1.816E-03 4.659E-02 2.105E-02 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 3.633E-03 9.319E-02 4.210E-02 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 1.653E-02 4.239E-01 1.915E-01 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 2.942E-02 7.547E-01 3.409E-01 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 5.422E-02 1.386E+00 6.259E-01 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 7.902E-02 2.016E+00 9.110E-01 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 9.187E-02 2.236E+00 1.012E+00 

0.000E+00 8.338E-05 1.008E-03 4.802E-04 1.047E-01 2.456E+00 1.113E+00 

1.000E+00 1.668E-04 2.016E-03 9.604E-04 1.445E-01 2.545E+00 1.172E+00 

2.000E+00 1.089E-03 1.141E-02 5.518E-03 1.843E-01 2.635E+00 1.230E+00 

3.000E+00 2.011E-03 2.081E-02 1.008E-02 3.056E-01 2.941E+00 1.425E+00 

4.000E+00 3.734E-03 2.748E-02 1.324E-02 4.269E-01 3.247E+00 1.620E+00 

5.000E+00 5.456E-03 3.415E-02 1.640E-02 5.362E-01 3.503E+00 1.793E+00 

6.000E+00 1.003E-02 7.461E-02 3.152E-02 6.455E-01 3.760E+00 1.966E+00 

7.000E+00 1.461E-02 1.151E-01 4.664E-02 9.045E-01 6.650E+00 2.922E+00 

8.000E+00 4.229E-02 1.779E-01 9.972E-02 1.163E+00 9.541E+00 3.878E+00 

9.000E+00 6.997E-02 2.407E-01 1.528E-01 4.001E+00 3.060E+01 1.236E+01 

1.000E+01 7.640E-02 2.632E-01 1.669E-01 6.839E+00 5.165E+01 2.085E+01 

1.100E+01 8.282E-02 2.857E-01 1.811E-01 2.188E+01 1.127E+02 5.578E+01 

1.200E+01 9.133E-02 3.313E-01 2.032E-01 3.692E+01 1.738E+02 9.071E+01 

1.300E+01 9.985E-02 3.769E-01 2.253E-01 6.715E+01 2.604E+02 1.538E+02 

1.400E+01 1.302E-01 6.143E-01 3.196E-01 9.738E+01 3.471E+02 2.168E+02 

1.500E+01 1.605E-01 8.517E-01 4.138E-01 1.170E+02 4.113E+02 2.592E+02 

1.600E+01 3.306E-01 1.305E+00 7.606E-01 1.367E+02 4.756E+02 3.015E+02 

1.700E+01 5.007E-01 1.759E+00 1.107E+00 1.388E+02 4.829E+02 3.061E+02 

1.800E+01 5.832E-01 2.037E+00 1.288E+00 1.409E+02 4.902E+02 3.107E+02 

1.900E+01 6.658E-01 2.316E+00 1.468E+00 1.436E+02 4.992E+02 3.164E+02 

2.000E+01 6.735E-01 2.343E+00 1.485E+00 1.463E+02 5.083E+02 3.221E+02 

2.100E+01 6.812E-01 2.369E+00 1.502E+00 1.523E+02 5.291E+02 3.353E+02 

2.200E+01 6.847E-01 2.382E+00 1.510E+00 1.583E+02 5.499E+02 3.486E+02 

2.300E+01 6.883E-01 2.395E+00 1.518E+00 1.646E+02 5.725E+02 3.629E+02 

2.400E+01 6.984E-01 2.427E+00 1.538E+00 1.710E+02 5.951E+02 3.772E+02 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-9 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

NOE3 NOE4 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 1.412E-04 3.621E-03 1.636E-03 

-4.000E+00 3.402E-02 8.725E-01 3.942E-01 1.412E-04 3.621E-03 1.636E-03 

-3.000E+00 1.369E-01 3.512E+00 1.587E+00 1.412E-04 3.621E-03 1.636E-03 

-2.000E+00 2.399E-01 6.152E+00 2.779E+00 1.412E-04 3.621E-03 1.636E-03 

-1.000E+00 5.787E-01 8.104E+00 3.811E+00 1.421E-03 3.645E-02 1.647E-02 

0.000E+00 9.176E-01 1.006E+01 4.843E+00 1.826E-03 4.604E-02 2.081E-02 

1.000E+00 2.986E+00 1.772E+01 9.158E+00 2.230E-03 5.562E-02 2.515E-02 

2.000E+00 5.055E+00 2.538E+01 1.347E+01 4.176E-03 9.278E-02 4.220E-02 

3.000E+00 8.539E+00 4.074E+01 2.153E+01 6.122E-03 1.299E-01 5.925E-02 

4.000E+00 1.202E+01 5.610E+01 2.960E+01 1.084E-02 1.533E-01 7.060E-02 

5.000E+00 2.164E+01 8.393E+01 4.940E+01 1.556E-02 1.766E-01 8.194E-02 

6.000E+00 3.125E+01 1.118E+02 6.920E+01 4.666E-02 3.673E-01 1.663E-01 

7.000E+00 4.053E+01 1.426E+02 8.923E+01 7.775E-02 5.580E-01 2.506E-01 

8.000E+00 4.980E+01 1.734E+02 1.093E+02 2.179E-01 1.157E+00 5.718E-01 

9.000E+00 6.013E+01 2.177E+02 1.336E+02 3.580E-01 1.755E+00 8.929E-01 

1.000E+01 7.045E+01 2.619E+02 1.580E+02 6.284E-01 2.425E+00 1.429E+00 

1.100E+01 9.705E+01 4.081E+02 2.282E+02 8.987E-01 3.094E+00 1.964E+00 

1.200E+01 1.236E+02 5.543E+02 2.983E+02 9.438E-01 3.251E+00 2.064E+00 

1.300E+01 2.255E+02 9.105E+02 5.242E+02 9.888E-01 3.408E+00 2.163E+00 

1.400E+01 3.273E+02 1.267E+03 7.502E+02 1.081E+00 4.033E+00 2.428E+00 

1.500E+01 4.597E+02 1.662E+03 1.027E+03 1.174E+00 4.657E+00 2.694E+00 

1.600E+01 5.921E+02 2.057E+03 1.304E+03 1.714E+00 7.458E+00 4.077E+00 

1.700E+01 6.265E+02 2.178E+03 1.381E+03 2.255E+00 1.026E+01 5.459E+00 

1.800E+01 6.608E+02 2.299E+03 1.457E+03 3.858E+00 1.467E+01 8.754E+00 

1.900E+01 6.730E+02 2.341E+03 1.484E+03 5.460E+00 1.908E+01 1.205E+01 

2.000E+01 6.851E+02 2.383E+03 1.511E+03 6.148E+00 2.143E+01 1.356E+01 

2.100E+01 7.032E+02 2.444E+03 1.549E+03 6.835E+00 2.377E+01 1.507E+01 

2.200E+01 7.213E+02 2.505E+03 1.588E+03 6.912E+00 2.403E+01 1.524E+01 

2.300E+01 7.537E+02 2.620E+03 1.660E+03 6.988E+00 2.430E+01 1.540E+01 

2.400E+01 7.861E+02 2.734E+03 1.733E+03 7.096E+00 2.466E+01 1.564E+01 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-10 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

NOE5 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 1.494E-01 3.832E+00 1.731E+00 

-8.000E+00 3.232E-01 8.289E+00 3.744E+00 

-7.000E+00 4.969E-01 1.275E+01 5.758E+00 

-6.000E+00 8.620E-01 2.163E+01 9.776E+00 

-5.000E+00 1.227E+00 3.052E+01 1.379E+01 

-4.000E+00 2.650E+00 4.470E+01 2.071E+01 

-3.000E+00 4.073E+00 5.888E+01 2.762E+01 

-2.000E+00 1.262E+01 9.192E+01 4.507E+01 

-1.000E+00 2.118E+01 1.250E+02 6.253E+01 

0.000E+00 5.627E+01 2.518E+02 1.382E+02 

1.000E+00 9.136E+01 3.786E+02 2.139E+02 

2.000E+00 1.762E+02 6.590E+02 3.968E+02 

3.000E+00 2.610E+02 9.394E+02 5.797E+02 

4.000E+00 3.461E+02 1.217E+03 7.623E+02 

5.000E+00 4.312E+02 1.496E+03 9.449E+02 

6.000E+00 5.048E+02 1.779E+03 1.114E+03 

7.000E+00 5.783E+02 2.062E+03 1.284E+03 

8.000E+00 7.841E+02 3.100E+03 1.806E+03 

9.000E+00 9.899E+02 4.138E+03 2.328E+03 

1.000E+01 1.638E+03 6.196E+03 3.719E+03 

1.100E+01 2.286E+03 8.255E+03 5.111E+03 

1.200E+01 2.946E+03 1.041E+04 6.528E+03 

1.300E+01 3.605E+03 1.256E+04 7.945E+03 

1.400E+01 3.990E+03 1.387E+04 8.786E+03 

1.500E+01 4.374E+03 1.519E+04 9.628E+03 

1.600E+01 4.604E+03 1.600E+04 1.014E+04 

1.700E+01 4.835E+03 1.681E+04 1.065E+04 

1.800E+01 4.999E+03 1.738E+04 1.101E+04 

1.900E+01 5.162E+03 1.794E+04 1.137E+04 

2.000E+01 5.354E+03 1.861E+04 1.180E+04 

2.100E+01 5.546E+03 1.929E+04 1.223E+04 

2.200E+01 5.676E+03 1.974E+04 1.252E+04 

2.300E+01 5.806E+03 2.020E+04 1.280E+04 

2.400E+01 5.891E+03 2.050E+04 1.299E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-11 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

OM1 OM2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 7.508E-01 1.315E+01 6.061E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 1.064E+00 1.680E+01 7.794E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 2.865E+00 4.129E+01 1.549E+01 

-9.000E+00 8.096E-02 2.986E-01 1.824E-01 4.666E+00 6.579E+01 2.318E+01 

-8.000E+00 1.619E-01 5.972E-01 3.649E-01 2.644E+01 1.332E+02 6.798E+01 

-7.000E+00 2.143E-01 8.507E-01 5.017E-01 4.822E+01 2.007E+02 1.128E+02 

-6.000E+00 2.668E-01 1.104E+00 6.385E-01 7.812E+01 2.879E+02 1.747E+02 

-5.000E+00 4.931E-01 3.167E+00 1.633E+00 1.080E+02 3.751E+02 2.367E+02 

-4.000E+00 7.194E-01 5.229E+00 2.627E+00 1.227E+02 4.244E+02 2.687E+02 

-3.000E+00 1.034E+00 1.004E+01 4.855E+00 1.374E+02 4.737E+02 3.006E+02 

-2.000E+00 1.348E+00 1.486E+01 7.084E+00 1.500E+02 5.207E+02 3.292E+02 

-1.000E+00 1.348E+00 1.486E+01 7.084E+00 1.627E+02 5.678E+02 3.578E+02 

0.000E+00 1.348E+00 1.486E+01 7.084E+00 2.053E+02 8.081E+02 4.717E+02 

1.000E+00 8.128E+00 5.889E+01 2.862E+01 2.480E+02 1.048E+03 5.856E+02 

2.000E+00 1.990E+01 1.046E+02 5.426E+01 4.147E+02 1.552E+03 9.372E+02 

3.000E+00 3.168E+01 1.504E+02 7.990E+01 5.814E+02 2.055E+03 1.289E+03 

4.000E+00 5.795E+01 2.323E+02 1.346E+02 7.098E+02 2.486E+03 1.567E+03 

5.000E+00 8.422E+01 3.142E+02 1.893E+02 8.382E+02 2.916E+03 1.845E+03 

6.000E+00 1.173E+02 4.216E+02 2.602E+02 9.228E+02 3.207E+03 2.032E+03 

7.000E+00 1.504E+02 5.290E+02 3.311E+02 1.007E+03 3.498E+03 2.218E+03 

8.000E+00 1.894E+02 6.735E+02 4.189E+02 1.057E+03 3.670E+03 2.327E+03 

9.000E+00 2.283E+02 8.179E+02 5.068E+02 1.106E+03 3.842E+03 2.436E+03 

1.000E+01 3.027E+02 1.115E+03 6.803E+02 1.162E+03 4.037E+03 2.560E+03 

1.100E+01 3.770E+02 1.412E+03 8.538E+02 1.218E+03 4.231E+03 2.683E+03 

1.200E+01 5.509E+02 2.030E+03 1.241E+03 1.295E+03 4.498E+03 2.852E+03 

1.300E+01 7.248E+02 2.648E+03 1.629E+03 1.372E+03 4.765E+03 3.021E+03 

1.400E+01 9.938E+02 3.563E+03 2.217E+03 1.444E+03 5.015E+03 3.179E+03 

1.500E+01 1.263E+03 4.479E+03 2.805E+03 1.515E+03 5.266E+03 3.338E+03 

1.600E+01 1.574E+03 5.527E+03 3.480E+03 1.563E+03 5.434E+03 3.445E+03 

1.700E+01 1.885E+03 6.576E+03 4.156E+03 1.610E+03 5.602E+03 3.551E+03 

1.800E+01 2.141E+03 7.452E+03 4.718E+03 1.631E+03 5.675E+03 3.598E+03 

1.900E+01 2.397E+03 8.327E+03 5.279E+03 1.652E+03 5.747E+03 3.644E+03 

2.000E+01 2.553E+03 8.868E+03 5.622E+03 1.671E+03 5.814E+03 3.686E+03 

2.100E+01 2.709E+03 9.409E+03 5.965E+03 1.690E+03 5.880E+03 3.727E+03 

2.200E+01 2.830E+03 9.834E+03 6.233E+03 1.708E+03 5.944E+03 3.768E+03 

2.300E+01 2.951E+03 1.026E+04 6.502E+03 1.727E+03 6.008E+03 3.809E+03 

2.400E+01 3.046E+03 1.059E+04 6.712E+03 1.745E+03 6.075E+03 3.851E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-12 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

OM3 OM4 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 1.007E-03 2.583E-02 1.167E-02 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 2.014E-03 5.166E-02 2.334E-02 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 1.252E-02 3.492E-01 1.214E-01 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 2.302E-02 6.467E-01 2.194E-01 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 2.178E-01 1.814E+00 7.472E-01 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 4.127E-01 2.982E+00 1.275E+00 

-5.000E+00 4.004E-02 5.220E-01 2.429E-01 1.170E+00 4.837E+00 2.746E+00 

-4.000E+00 8.009E-02 1.044E+00 4.858E-01 1.928E+00 6.692E+00 4.216E+00 

-3.000E+00 5.456E-01 5.165E+00 2.499E+00 2.362E+00 8.299E+00 5.197E+00 

-2.000E+00 1.011E+00 9.287E+00 4.511E+00 2.797E+00 9.907E+00 6.177E+00 

-1.000E+00 1.907E+00 2.050E+01 9.799E+00 3.471E+00 1.327E+01 7.884E+00 

0.000E+00 2.802E+00 3.172E+01 1.509E+01 4.145E+00 1.663E+01 9.590E+00 

1.000E+00 5.188E+00 4.896E+01 2.373E+01 6.205E+00 2.702E+01 1.478E+01 

2.000E+00 7.574E+00 6.621E+01 3.237E+01 8.265E+00 3.740E+01 1.996E+01 

3.000E+00 1.976E+01 1.280E+02 6.116E+01 1.507E+01 6.099E+01 3.488E+01 

4.000E+00 3.194E+01 1.899E+02 8.995E+01 2.187E+01 8.458E+01 4.979E+01 

5.000E+00 7.468E+01 3.231E+02 1.791E+02 3.390E+01 1.232E+02 7.557E+01 

6.000E+00 1.174E+02 4.563E+02 2.683E+02 4.593E+01 1.618E+02 1.013E+02 

7.000E+00 1.748E+02 6.341E+02 3.891E+02 5.789E+01 2.034E+02 1.276E+02 

8.000E+00 2.323E+02 8.119E+02 5.100E+02 6.985E+01 2.449E+02 1.539E+02 

9.000E+00 2.809E+02 9.771E+02 6.158E+02 8.648E+01 3.039E+02 1.908E+02 

1.000E+01 3.295E+02 1.142E+03 7.216E+02 1.031E+02 3.630E+02 2.277E+02 

1.100E+01 3.970E+02 1.450E+03 8.881E+02 1.294E+02 4.581E+02 2.866E+02 

1.200E+01 4.645E+02 1.757E+03 1.055E+03 1.557E+02 5.533E+02 3.455E+02 

1.300E+01 6.922E+02 2.586E+03 1.567E+03 1.980E+02 6.962E+02 4.376E+02 

1.400E+01 9.199E+02 3.415E+03 2.079E+03 2.403E+02 8.391E+02 5.296E+02 

1.500E+01 1.308E+03 4.718E+03 2.923E+03 2.847E+02 9.930E+02 6.271E+02 

1.600E+01 1.696E+03 6.020E+03 3.767E+03 3.291E+02 1.147E+03 7.247E+02 

1.700E+01 2.119E+03 7.431E+03 4.680E+03 3.807E+02 1.325E+03 8.379E+02 

1.800E+01 2.542E+03 8.842E+03 5.593E+03 4.323E+02 1.503E+03 9.512E+02 

1.900E+01 2.867E+03 9.965E+03 6.309E+03 4.849E+02 1.685E+03 1.067E+03 

2.000E+01 3.192E+03 1.109E+04 7.025E+03 5.375E+02 1.868E+03 1.183E+03 

2.100E+01 3.467E+03 1.204E+04 7.632E+03 5.827E+02 2.024E+03 1.283E+03 

2.200E+01 3.742E+03 1.300E+04 8.239E+03 6.279E+02 2.180E+03 1.382E+03 

2.300E+01 3.986E+03 1.385E+04 8.778E+03 6.567E+02 2.282E+03 1.446E+03 

2.400E+01 4.231E+03 1.469E+04 9.316E+03 6.855E+02 2.383E+03 1.510E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-13 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

OM5 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 8.259E-01 3.053E+00 1.861E+00 

-3.000E+00 1.652E+00 6.106E+00 3.721E+00 

-2.000E+00 3.532E+00 2.417E+01 1.235E+01 

-1.000E+00 5.412E+00 4.222E+01 2.097E+01 

0.000E+00 7.129E+00 6.399E+01 3.118E+01 

1.000E+00 8.845E+00 8.575E+01 4.140E+01 

2.000E+00 1.223E+01 1.145E+02 5.566E+01 

3.000E+00 1.562E+01 1.433E+02 6.992E+01 

4.000E+00 3.979E+01 2.925E+02 1.325E+02 

5.000E+00 6.397E+01 4.417E+02 1.950E+02 

6.000E+00 1.752E+02 7.693E+02 4.227E+02 

7.000E+00 2.865E+02 1.097E+03 6.504E+02 

8.000E+00 4.201E+02 1.520E+03 9.339E+02 

9.000E+00 5.537E+02 1.943E+03 1.217E+03 

1.000E+01 6.791E+02 2.364E+03 1.490E+03 

1.100E+01 8.045E+02 2.786E+03 1.762E+03 

1.200E+01 9.419E+02 3.305E+03 2.075E+03 

1.300E+01 1.079E+03 3.825E+03 2.389E+03 

1.400E+01 1.436E+03 5.444E+03 3.262E+03 

1.500E+01 1.794E+03 7.063E+03 4.134E+03 

1.600E+01 2.786E+03 1.028E+04 6.275E+03 

1.700E+01 3.777E+03 1.350E+04 8.415E+03 

1.800E+01 4.850E+03 1.707E+04 1.073E+04 

1.900E+01 5.923E+03 2.064E+04 1.305E+04 

2.000E+01 6.808E+03 2.369E+04 1.499E+04 

2.100E+01 7.694E+03 2.673E+04 1.694E+04 

2.200E+01 8.312E+03 2.888E+04 1.831E+04 

2.300E+01 8.930E+03 3.102E+04 1.967E+04 

2.400E+01 9.564E+03 3.322E+04 2.106E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-14 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

SB1 SB3 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

1.000E+00 1.226E+00 2.429E+01 1.111E+01 5.927E-01 4.089E+00 2.070E+00 

2.000E+00 2.453E+00 4.859E+01 2.222E+01 1.185E+00 8.178E+00 4.141E+00 

3.000E+00 2.597E+00 4.999E+01 2.290E+01 3.716E+00 2.489E+01 1.267E+01 

4.000E+00 2.742E+00 5.139E+01 2.358E+01 6.246E+00 4.159E+01 2.120E+01 

5.000E+00 6.060E+00 1.307E+02 4.488E+01 7.259E+00 4.414E+01 2.292E+01 

6.000E+00 9.377E+00 2.099E+02 6.618E+01 8.272E+00 4.668E+01 2.464E+01 

7.000E+00 8.578E+01 4.743E+02 2.294E+02 3.417E+01 1.340E+02 8.054E+01 

8.000E+00 1.622E+02 7.387E+02 3.926E+02 6.006E+01 2.213E+02 1.364E+02 

9.000E+00 2.826E+02 1.072E+03 6.403E+02 1.576E+02 5.524E+02 3.486E+02 

1.000E+01 4.031E+02 1.405E+03 8.879E+02 2.552E+02 8.836E+02 5.607E+02 

1.100E+01 4.096E+02 1.424E+03 9.015E+02 2.552E+02 8.836E+02 5.607E+02 

1.200E+01 4.162E+02 1.442E+03 9.151E+02 2.552E+02 8.836E+02 5.607E+02 

1.300E+01 4.162E+02 1.442E+03 9.151E+02 2.624E+02 1.040E+03 6.035E+02 

1.400E+01 4.162E+02 1.442E+03 9.152E+02 2.696E+02 1.197E+03 6.462E+02 

1.500E+01 4.586E+02 2.103E+03 1.114E+03 4.215E+02 2.120E+03 1.061E+03 

1.600E+01 5.009E+02 2.764E+03 1.312E+03 5.735E+02 3.043E+03 1.477E+03 

1.700E+01 1.117E+03 4.493E+03 2.590E+03 1.077E+03 4.266E+03 2.479E+03 

1.800E+01 1.733E+03 6.223E+03 3.867E+03 1.580E+03 5.490E+03 3.481E+03 

1.900E+01 2.161E+03 7.627E+03 4.793E+03 1.624E+03 5.649E+03 3.580E+03 

2.000E+01 2.590E+03 9.031E+03 5.718E+03 1.667E+03 5.808E+03 3.679E+03 

2.100E+01 2.611E+03 9.101E+03 5.764E+03 1.667E+03 5.808E+03 3.679E+03 

2.200E+01 2.632E+03 9.170E+03 5.809E+03 1.667E+03 5.808E+03 3.680E+03 

2.300E+01 2.656E+03 9.237E+03 5.854E+03 1.675E+03 5.833E+03 3.695E+03 

2.400E+01 2.679E+03 9.304E+03 5.899E+03 1.683E+03 5.859E+03 3.711E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-15 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

SB4 SB5 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

1.000E+00 1.997E-02 3.719E-01 1.711E-01 0.000E+00 0.000E+00 0.000E+00 

2.000E+00 3.994E-02 7.439E-01 3.423E-01 0.000E+00 0.000E+00 0.000E+00 

3.000E+00 2.766E-01 4.299E+00 1.992E+00 1.666E-05 6.477E-05 3.834E-05 

4.000E+00 5.134E-01 7.855E+00 3.641E+00 3.333E-05 1.295E-04 7.668E-05 

5.000E+00 1.277E+00 1.081E+01 5.174E+00 2.959E-03 7.356E-02 3.327E-02 

6.000E+00 2.041E+00 1.376E+01 6.707E+00 5.884E-03 1.470E-01 6.647E-02 

7.000E+00 4.988E+00 2.563E+01 1.312E+01 1.039E-02 1.924E-01 8.391E-02 

8.000E+00 7.936E+00 3.749E+01 1.954E+01 1.490E-02 2.378E-01 1.014E-01 

9.000E+00 1.495E+01 5.921E+01 3.437E+01 5.916E-01 2.146E+00 1.336E+00 

1.000E+01 2.197E+01 8.092E+01 4.921E+01 1.168E+00 4.055E+00 2.571E+00 

1.100E+01 2.976E+01 1.054E+02 6.568E+01 1.199E+00 4.157E+00 2.637E+00 

1.200E+01 3.756E+01 1.298E+02 8.215E+01 1.229E+00 4.259E+00 2.703E+00 

1.300E+01 4.235E+01 1.550E+02 9.445E+01 1.229E+00 4.259E+00 2.703E+00 

1.400E+01 4.715E+01 1.801E+02 1.068E+02 1.229E+00 4.259E+00 2.703E+00 

1.500E+01 4.965E+01 2.663E+02 1.282E+02 1.230E+00 4.260E+00 2.703E+00 

1.600E+01 5.216E+01 3.525E+02 1.497E+02 1.230E+00 4.260E+00 2.703E+00 

1.700E+01 6.059E+01 3.733E+02 1.666E+02 1.405E+00 8.264E+00 3.781E+00 

1.800E+01 6.902E+01 3.941E+02 1.834E+02 1.580E+00 1.227E+01 4.859E+00 

1.900E+01 1.457E+02 5.831E+02 3.365E+02 4.637E+00 1.971E+01 1.096E+01 

2.000E+01 2.224E+02 7.721E+02 4.896E+02 7.694E+00 2.716E+01 1.705E+01 

2.100E+01 2.299E+02 7.993E+02 5.066E+02 7.854E+00 2.754E+01 1.737E+01 

2.200E+01 2.374E+02 8.265E+02 5.237E+02 8.014E+00 2.792E+01 1.769E+01 

2.300E+01 2.381E+02 8.288E+02 5.251E+02 8.014E+00 2.792E+01 1.769E+01 

2.400E+01 2.388E+02 8.311E+02 5.265E+02 8.014E+00 2.792E+01 1.769E+01 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-16 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

JE1 JE2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 5.896E-02 1.512E+00 6.831E-01 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 6.633E-02 1.701E+00 7.685E-01 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 7.369E-02 1.890E+00 8.539E-01 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 5.672E-01 6.410E+00 2.458E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 1.061E+00 1.093E+01 4.062E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 5.702E+00 2.459E+01 1.362E+01 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 1.034E+01 3.825E+01 2.317E+01 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 1.636E+01 5.973E+01 3.658E+01 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 2.237E+01 8.121E+01 4.998E+01 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 2.781E+01 1.045E+02 6.323E+01 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 3.326E+01 1.278E+02 7.647E+01 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 4.322E+01 2.323E+02 1.129E+02 

1.000E+00 0.000E+00 0.000E+00 0.000E+00 5.319E+01 3.368E+02 1.493E+02 

2.000E+00 0.000E+00 0.000E+00 0.000E+00 1.432E+02 6.527E+02 3.477E+02 

3.000E+00 1.666E-05 6.477E-05 3.834E-05 2.332E+02 9.686E+02 5.461E+02 

4.000E+00 3.333E-05 1.295E-04 7.668E-05 3.702E+02 1.380E+03 8.331E+02 

5.000E+00 2.959E-03 7.356E-02 3.327E-02 5.073E+02 1.792E+03 1.120E+03 

6.000E+00 5.884E-03 1.470E-01 6.647E-02 6.136E+02 2.142E+03 1.350E+03 

7.000E+00 1.039E-02 1.924E-01 8.391E-02 7.200E+02 2.491E+03 1.579E+03 

8.000E+00 1.490E-02 2.378E-01 1.014E-01 7.843E+02 2.715E+03 1.720E+03 

9.000E+00 5.916E-01 2.146E+00 1.336E+00 8.486E+02 2.939E+03 1.861E+03 

1.000E+01 1.168E+00 4.055E+00 2.571E+00 9.899E+02 3.639E+03 2.223E+03 

1.100E+01 1.199E+00 4.157E+00 2.637E+00 1.131E+03 4.339E+03 2.584E+03 

1.200E+01 1.229E+00 4.259E+00 2.703E+00 1.658E+03 6.045E+03 3.716E+03 

1.300E+01 1.229E+00 4.259E+00 2.703E+00 2.185E+03 7.752E+03 4.848E+03 

1.400E+01 1.229E+00 4.259E+00 2.703E+00 2.664E+03 9.342E+03 5.886E+03 

1.500E+01 1.230E+00 4.260E+00 2.703E+00 3.143E+03 1.093E+04 6.924E+03 

1.600E+01 1.230E+00 4.260E+00 2.703E+00 3.416E+03 1.188E+04 7.524E+03 

1.700E+01 1.405E+00 8.264E+00 3.781E+00 3.688E+03 1.282E+04 8.124E+03 

1.800E+01 1.580E+00 1.227E+01 4.859E+00 3.845E+03 1.336E+04 8.469E+03 

1.900E+01 4.637E+00 1.971E+01 1.096E+01 4.001E+03 1.390E+04 8.813E+03 

2.000E+01 7.694E+00 2.716E+01 1.705E+01 4.164E+03 1.447E+04 9.171E+03 

2.100E+01 7.854E+00 2.754E+01 1.737E+01 4.327E+03 1.503E+04 9.529E+03 

2.200E+01 8.014E+00 2.792E+01 1.769E+01 4.504E+03 1.565E+04 9.923E+03 

2.300E+01 8.014E+00 2.792E+01 1.769E+01 4.680E+03 1.627E+04 1.032E+04 

2.400E+01 8.014E+00 2.792E+01 1.769E+01 4.779E+03 1.662E+04 1.054E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-17 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

JE3 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 4.069E-02 7.390E-01 3.409E-01 

-8.000E+00 8.138E-02 1.478E+00 6.818E-01 

-7.000E+00 1.340E+00 1.135E+01 4.916E+00 

-6.000E+00 2.599E+00 2.122E+01 9.151E+00 

-5.000E+00 1.248E+01 5.682E+01 3.127E+01 

-4.000E+00 2.235E+01 9.243E+01 5.338E+01 

-3.000E+00 3.686E+01 1.641E+02 9.256E+01 

-2.000E+00 5.137E+01 2.358E+02 1.317E+02 

-1.000E+00 6.634E+01 3.098E+02 1.691E+02 

0.000E+00 8.132E+01 3.838E+02 2.065E+02 

1.000E+00 1.568E+02 8.388E+02 4.084E+02 

2.000E+00 2.322E+02 1.294E+03 6.102E+02 

3.000E+00 4.716E+02 1.963E+03 1.105E+03 

4.000E+00 7.110E+02 2.631E+03 1.599E+03 

5.000E+00 9.994E+02 3.541E+03 2.210E+03 

6.000E+00 1.288E+03 4.450E+03 2.821E+03 

7.000E+00 1.380E+03 4.775E+03 3.026E+03 

8.000E+00 1.472E+03 5.100E+03 3.230E+03 

9.000E+00 1.715E+03 6.406E+03 3.874E+03 

1.000E+01 1.958E+03 7.712E+03 4.518E+03 

1.100E+01 3.102E+03 1.189E+04 7.098E+03 

1.200E+01 4.246E+03 1.607E+04 9.678E+03 

1.300E+01 6.014E+03 2.155E+04 1.341E+04 

1.400E+01 7.782E+03 2.703E+04 1.714E+04 

1.500E+01 8.261E+03 2.872E+04 1.820E+04 

1.600E+01 8.740E+03 3.040E+04 1.927E+04 

1.700E+01 8.931E+03 3.106E+04 1.969E+04 

1.800E+01 9.121E+03 3.173E+04 2.011E+04 

1.900E+01 9.362E+03 3.254E+04 2.063E+04 

2.000E+01 9.602E+03 3.335E+04 2.114E+04 

2.100E+01 1.008E+04 3.504E+04 2.221E+04 

2.200E+01 1.056E+04 3.673E+04 2.328E+04 

2.300E+01 1.085E+04 3.773E+04 2.391E+04 

2.400E+01 1.113E+04 3.872E+04 2.455E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-18 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

JW1 JW2 

Elevation (ft) 5% 95% Mean 5% 25% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 8.314E-03 1.987E-01 8.998E-02 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 1.663E-02 3.974E-01 1.800E-01 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 2.193E-02 4.154E-01 1.903E-01 4.000E-05 1.936E-04 4.635E-04 

-5.000E+00 2.722E-02 4.334E-01 2.007E-01 8.000E-05 3.872E-04 9.270E-04 

-4.000E+00 9.473E-02 1.509E+00 6.987E-01 2.157E-03 9.819E-03 2.334E-02 

-3.000E+00 1.622E-01 2.584E+00 1.197E+00 4.233E-03 1.925E-02 4.574E-02 

-2.000E+00 2.833E-01 3.087E+00 1.476E+00 3.649E-02 1.518E-01 3.557E-01 

-1.000E+00 4.044E-01 3.589E+00 1.754E+00 6.875E-02 2.843E-01 6.656E-01 

0.000E+00 1.151E+00 8.715E+00 3.769E+00 2.192E-01 7.510E-01 1.694E+00 

1.000E+00 1.897E+00 1.384E+01 5.783E+00 3.696E-01 1.218E+00 2.722E+00 

2.000E+00 5.192E+00 2.445E+01 1.278E+01 6.380E-01 1.754E+00 3.725E+00 

3.000E+00 8.486E+00 3.506E+01 1.978E+01 9.063E-01 2.290E+00 4.728E+00 

4.000E+00 1.410E+01 5.164E+01 3.146E+01 1.538E+00 3.342E+00 6.310E+00 

5.000E+00 1.971E+01 6.821E+01 4.315E+01 2.169E+00 4.395E+00 7.893E+00 

6.000E+00 2.150E+01 7.447E+01 4.706E+01 5.709E+00 9.938E+00 1.586E+01 

7.000E+00 2.330E+01 8.072E+01 5.097E+01 9.249E+00 1.548E+01 2.384E+01 

8.000E+00 2.795E+01 1.052E+02 6.298E+01 1.742E+01 2.873E+01 4.037E+01 

9.000E+00 3.260E+01 1.296E+02 7.499E+01 2.560E+01 4.198E+01 5.691E+01 

1.000E+01 5.370E+01 2.170E+02 1.244E+02 3.374E+01 5.531E+01 7.424E+01 

1.100E+01 7.481E+01 3.044E+02 1.739E+02 4.187E+01 6.864E+01 9.156E+01 

1.200E+01 1.176E+02 4.309E+02 2.631E+02 4.548E+01 7.461E+01 1.002E+02 

1.300E+01 1.605E+02 5.575E+02 3.524E+02 4.908E+01 8.058E+01 1.088E+02 

1.400E+01 1.755E+02 6.083E+02 3.853E+02 6.463E+01 1.077E+02 1.555E+02 

1.500E+01 1.905E+02 6.591E+02 4.183E+02 8.019E+01 1.348E+02 2.021E+02 

1.600E+01 2.026E+02 7.021E+02 4.448E+02 1.652E+02 2.746E+02 3.846E+02 

1.700E+01 2.147E+02 7.451E+02 4.713E+02 2.503E+02 4.144E+02 5.671E+02 

1.800E+01 2.517E+02 8.863E+02 5.567E+02 3.424E+02 5.656E+02 7.617E+02 

1.900E+01 2.888E+02 1.028E+03 6.421E+02 4.345E+02 7.167E+02 9.563E+02 

2.000E+01 3.675E+02 1.291E+03 8.120E+02 4.547E+02 7.506E+02 1.002E+03 

2.100E+01 4.462E+02 1.554E+03 9.820E+02 4.749E+02 7.844E+02 1.048E+03 

2.200E+01 5.117E+02 1.779E+03 1.126E+03 4.803E+02 7.935E+02 1.059E+03 

2.300E+01 5.772E+02 2.003E+03 1.270E+03 4.858E+02 8.026E+02 1.071E+03 

2.400E+01 6.138E+02 2.132E+03 1.351E+03 4.989E+02 8.235E+02 1.099E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-19 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

JW3 JW4 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 4.917E-04 7.954E-03 3.705E-03 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 9.834E-04 1.591E-02 7.410E-03 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 1.237E-02 1.622E-01 7.666E-02 

-7.000E+00 3.717E-04 9.533E-03 4.306E-03 2.376E-02 3.086E-01 1.459E-01 

-6.000E+00 7.433E-04 1.907E-02 8.613E-03 1.144E-01 1.235E+00 5.928E-01 

-5.000E+00 9.020E-03 1.982E-01 7.939E-02 2.050E-01 2.160E+00 1.040E+00 

-4.000E+00 1.730E-02 3.774E-01 1.502E-01 6.534E-01 5.894E+00 2.883E+00 

-3.000E+00 1.981E-01 1.797E+00 7.772E-01 1.102E+00 9.628E+00 4.727E+00 

-2.000E+00 3.789E-01 3.217E+00 1.404E+00 3.240E+00 2.481E+01 1.240E+01 

-1.000E+00 1.843E+00 9.844E+00 5.188E+00 5.377E+00 4.000E+01 2.006E+01 

0.000E+00 3.307E+00 1.647E+01 8.973E+00 9.448E+00 6.187E+01 3.172E+01 

1.000E+00 5.694E+00 3.084E+01 1.644E+01 1.352E+01 8.373E+01 4.337E+01 

2.000E+00 8.081E+00 4.520E+01 2.390E+01 2.237E+01 1.188E+02 6.325E+01 

3.000E+00 1.092E+01 6.322E+01 3.313E+01 3.122E+01 1.539E+02 8.313E+01 

4.000E+00 1.376E+01 8.123E+01 4.237E+01 5.328E+01 2.417E+02 1.322E+02 

5.000E+00 2.080E+01 1.179E+02 6.145E+01 7.534E+01 3.294E+02 1.813E+02 

6.000E+00 2.785E+01 1.547E+02 8.053E+01 1.292E+02 5.222E+02 2.997E+02 

7.000E+00 5.421E+01 2.563E+02 1.378E+02 1.830E+02 7.149E+02 4.181E+02 

8.000E+00 8.058E+01 3.580E+02 1.951E+02 2.729E+02 9.888E+02 6.066E+02 

9.000E+00 1.381E+02 5.400E+02 3.166E+02 3.627E+02 1.263E+03 7.952E+02 

1.000E+01 1.957E+02 7.221E+02 4.381E+02 4.401E+02 1.547E+03 9.695E+02 

1.100E+01 2.709E+02 9.742E+02 6.016E+02 5.174E+02 1.831E+03 1.144E+03 

1.200E+01 3.461E+02 1.226E+03 7.650E+02 6.727E+02 2.430E+03 1.500E+03 

1.300E+01 4.495E+02 1.604E+03 9.975E+02 8.281E+02 3.028E+03 1.856E+03 

1.400E+01 5.529E+02 1.982E+03 1.230E+03 1.167E+03 4.479E+03 2.663E+03 

1.500E+01 7.401E+02 2.653E+03 1.647E+03 1.506E+03 5.929E+03 3.470E+03 

1.600E+01 9.274E+02 3.323E+03 2.064E+03 2.267E+03 8.265E+03 5.078E+03 

1.700E+01 1.243E+03 4.558E+03 2.793E+03 3.029E+03 1.060E+04 6.687E+03 

1.800E+01 1.558E+03 5.794E+03 3.523E+03 3.575E+03 1.246E+04 7.877E+03 

1.900E+01 2.189E+03 7.862E+03 4.883E+03 4.121E+03 1.432E+04 9.067E+03 

2.000E+01 2.821E+03 9.930E+03 6.243E+03 4.552E+03 1.581E+04 1.002E+04 

2.100E+01 3.415E+03 1.194E+04 7.534E+03 4.982E+03 1.729E+04 1.096E+04 

2.200E+01 4.009E+03 1.395E+04 8.825E+03 5.255E+03 1.824E+04 1.156E+04 

2.300E+01 4.477E+03 1.556E+04 9.856E+03 5.528E+03 1.919E+04 1.217E+04 

2.400E+01 4.945E+03 1.717E+04 1.089E+04 5.821E+03 2.022E+04 1.282E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-20 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Continued) 

PL11 

Elevation (ft) 5% 95% Mean 

-1.200E+01 2.767E-01 1.212E+00 6.872E-01 

-1.100E+01 3.432E-01 1.776E+00 9.593E-01 

-1.000E+01 4.097E-01 2.341E+00 1.231E+00 

-9.000E+00 4.555E-01 2.541E+00 1.345E+00 

-8.000E+00 5.013E-01 2.740E+00 1.458E+00 

-7.000E+00 7.744E-01 6.113E+00 2.544E+00 

-6.000E+00 1.048E+00 9.487E+00 3.631E+00 

-5.000E+00 4.072E+00 1.803E+01 9.811E+00 

-4.000E+00 7.097E+00 2.657E+01 1.599E+01 

-3.000E+00 1.002E+01 3.558E+01 2.214E+01 

-2.000E+00 1.294E+01 4.459E+01 2.829E+01 

-1.000E+00 1.413E+01 4.865E+01 3.089E+01 

0.000E+00 1.532E+01 5.272E+01 3.349E+01 

1.000E+00 1.690E+01 5.953E+01 3.729E+01 

2.000E+00 1.848E+01 6.633E+01 4.109E+01 

3.000E+00 2.597E+01 1.064E+02 6.045E+01 

4.000E+00 3.346E+01 1.465E+02 7.980E+01 

5.000E+00 5.770E+01 2.152E+02 1.298E+02 

6.000E+00 8.193E+01 2.839E+02 1.798E+02 

7.000E+00 9.195E+01 3.184E+02 2.017E+02 

8.000E+00 1.020E+02 3.530E+02 2.237E+02 

9.000E+00 1.127E+02 3.905E+02 2.474E+02 

1.000E+01 1.235E+02 4.280E+02 2.710E+02 

1.100E+01 1.384E+02 4.797E+02 3.037E+02 

1.200E+01 1.533E+02 5.314E+02 3.364E+02 

1.300E+01 1.744E+02 6.062E+02 3.833E+02 

1.400E+01 1.955E+02 6.809E+02 4.301E+02 

1.500E+01 2.253E+02 7.829E+02 4.956E+02 

1.600E+01 2.551E+02 8.849E+02 5.611E+02 

1.700E+01 2.709E+02 9.403E+02 5.959E+02 

1.800E+01 2.868E+02 9.957E+02 6.307E+02 

1.900E+01 3.195E+02 1.113E+03 7.038E+02 

2.000E+01 3.523E+02 1.231E+03 7.770E+02 

2.100E+01 4.088E+02 1.424E+03 9.008E+02 

2.200E+01 4.653E+02 1.616E+03 1.025E+03 

2.300E+01 4.964E+02 1.724E+03 1.093E+03 

2.400E+01 5.275E+02 1.832E+03 1.162E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-21 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-3 Pre-Katrina – Stage-Fatality (Concluded) 

SC1 SC2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

2.000E+00 0.000E+00 0.000E+00 0.000E+00 3.671E-03 9.415E-02 4.253E-02 

3.000E+00 0.000E+00 0.000E+00 0.000E+00 7.341E-03 1.883E-01 8.506E-02 

4.000E+00 0.000E+00 0.000E+00 0.000E+00 5.968E-02 1.521E+00 6.874E-01 

5.000E+00 0.000E+00 0.000E+00 0.000E+00 1.120E-01 2.854E+00 1.290E+00 

6.000E+00 9.412E-06 2.414E-04 1.091E-04 4.273E-01 4.618E+00 2.210E+00 

7.000E+00 1.882E-05 4.828E-04 2.181E-04 7.426E-01 6.381E+00 3.130E+00 

8.000E+00 1.604E-02 5.968E-02 3.687E-02 1.863E+00 1.486E+01 6.670E+00 

9.000E+00 3.207E-02 1.189E-01 7.353E-02 2.984E+00 2.334E+01 1.021E+01 

1.000E+01 3.867E-01 1.340E+00 8.474E-01 9.355E+00 4.333E+01 2.330E+01 

1.100E+01 7.413E-01 2.561E+00 1.621E+00 1.573E+01 6.332E+01 3.638E+01 

1.200E+01 1.766E+00 6.074E+00 3.856E+00 2.634E+01 1.005E+02 5.967E+01 

1.300E+01 2.790E+00 9.587E+00 6.091E+00 3.695E+01 1.377E+02 8.296E+01 

1.400E+01 2.940E+00 1.011E+01 6.419E+00 5.326E+01 1.896E+02 1.176E+02 

1.500E+01 3.089E+00 1.063E+01 6.748E+00 6.957E+01 2.415E+02 1.523E+02 

1.600E+01 3.298E+00 1.233E+01 7.408E+00 8.632E+01 3.043E+02 1.904E+02 

1.700E+01 3.506E+00 1.404E+01 8.068E+00 1.031E+02 3.670E+02 2.285E+02 

1.800E+01 5.428E+00 3.787E+01 1.583E+01 1.462E+02 5.658E+02 3.341E+02 

1.900E+01 7.349E+00 6.170E+01 2.359E+01 1.893E+02 7.645E+02 4.398E+02 

2.000E+01 2.279E+01 9.731E+01 5.392E+01 3.020E+02 1.115E+03 6.796E+02 

2.100E+01 3.823E+01 1.329E+02 8.425E+01 4.147E+02 1.466E+03 9.194E+02 

2.200E+01 3.917E+01 1.362E+02 8.636E+01 5.214E+02 1.840E+03 1.155E+03 

2.300E+01 4.011E+01 1.396E+02 8.848E+01 6.281E+02 2.215E+03 1.391E+03 

2.400E+01 4.044E+01 1.407E+02 8.918E+01 7.713E+02 2.696E+03 1.701E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-22 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality 

OW1 OW2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 8.780E-03 2.252E-01 1.017E-01 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 1.756E-02 4.504E-01 2.035E-01 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 1.201E-01 1.960E+00 8.930E-01 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 2.227E-01 3.469E+00 1.583E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 1.190E+00 7.069E+00 3.710E+00 

-3.000E+00 5.647E-05 1.448E-03 6.543E-04 2.158E+00 1.067E+01 5.838E+00 

-2.000E+00 1.129E-04 2.897E-03 1.309E-03 3.304E+00 1.828E+01 9.716E+00 

-1.000E+00 4.753E-04 1.219E-02 5.507E-03 4.449E+00 2.590E+01 1.359E+01 

0.000E+00 8.376E-04 2.149E-02 9.705E-03 6.732E+00 3.512E+01 1.890E+01 

1.000E+00 1.176E-03 3.018E-02 1.363E-02 9.015E+00 4.435E+01 2.420E+01 

2.000E+00 1.515E-03 3.887E-02 1.756E-02 1.444E+01 7.418E+01 3.821E+01 

3.000E+00 5.990E-03 1.442E-01 6.528E-02 1.987E+01 1.040E+02 5.222E+01 

4.000E+00 1.046E-02 2.495E-01 1.130E-01 4.191E+01 1.896E+02 1.019E+02 

5.000E+00 3.548E-02 7.692E-01 3.500E-01 6.394E+01 2.753E+02 1.515E+02 

6.000E+00 6.049E-02 1.289E+00 5.870E-01 1.091E+02 4.107E+02 2.458E+02 

7.000E+00 8.772E-02 1.410E+00 6.413E-01 1.542E+02 5.461E+02 3.401E+02 

8.000E+00 1.150E-01 1.531E+00 6.956E-01 1.929E+02 6.767E+02 4.242E+02 

9.000E+00 2.917E-01 2.555E+00 1.149E+00 2.315E+02 8.073E+02 5.083E+02 

1.000E+01 4.685E-01 3.579E+00 1.602E+00 2.847E+02 1.007E+03 6.292E+02 

1.100E+01 1.375E+00 6.980E+00 3.554E+00 3.380E+02 1.206E+03 7.502E+02 

1.200E+01 2.281E+00 1.038E+01 5.506E+00 4.542E+02 1.686E+03 1.025E+03 

1.300E+01 3.838E+00 1.448E+01 8.647E+00 5.705E+02 2.165E+03 1.299E+03 

1.400E+01 5.395E+00 1.857E+01 1.179E+01 8.353E+02 3.025E+03 1.867E+03 

1.500E+01 5.829E+00 2.009E+01 1.275E+01 1.100E+03 3.886E+03 2.436E+03 

1.600E+01 6.263E+00 2.161E+01 1.371E+01 1.350E+03 4.722E+03 2.976E+03 

1.700E+01 7.047E+00 2.789E+01 1.616E+01 1.599E+03 5.559E+03 3.516E+03 

1.800E+01 7.832E+00 3.416E+01 1.860E+01 1.823E+03 6.333E+03 4.008E+03 

1.900E+01 1.322E+01 6.911E+01 3.380E+01 2.046E+03 7.107E+03 4.499E+03 

2.000E+01 1.862E+01 1.041E+02 4.900E+01 2.293E+03 7.967E+03 5.043E+03 

2.100E+01 3.827E+01 1.527E+02 8.827E+01 2.539E+03 8.828E+03 5.588E+03 

2.200E+01 5.792E+01 2.013E+02 1.275E+02 2.804E+03 9.747E+03 6.173E+03 

2.300E+01 6.193E+01 2.153E+02 1.365E+02 3.070E+03 1.067E+04 6.758E+03 

2.400E+01 6.594E+01 2.293E+02 1.454E+02 3.305E+03 1.149E+04 7.278E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-23 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

NOE1 NOE2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 8.941E-05 2.293E-03 1.036E-03 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 1.788E-04 4.587E-03 2.072E-03 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 8.188E-04 2.100E-02 9.488E-03 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 1.459E-03 3.742E-02 1.690E-02 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 2.717E-03 6.908E-02 3.121E-02 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 3.975E-03 1.007E-01 4.553E-02 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 4.883E-03 1.124E-01 5.103E-02 

0.000E+00 2.353E-06 6.035E-05 2.726E-05 5.792E-03 1.240E-01 5.653E-02 

1.000E+00 4.706E-06 1.207E-04 5.453E-05 1.014E-02 1.374E-01 6.438E-02 

2.000E+00 2.588E-05 6.639E-04 2.999E-04 1.449E-02 1.508E-01 7.223E-02 

3.000E+00 4.706E-05 1.207E-03 5.453E-04 3.319E-02 2.091E-01 1.077E-01 

4.000E+00 4.835E-05 1.210E-03 5.473E-04 5.190E-02 2.673E-01 1.433E-01 

5.000E+00 4.963E-05 1.213E-03 5.494E-04 8.603E-02 3.542E-01 2.065E-01 

6.000E+00 7.061E-05 1.256E-03 5.799E-04 1.202E-01 4.411E-01 2.697E-01 

7.000E+00 9.158E-05 1.299E-03 6.104E-04 1.459E-01 6.190E-01 3.436E-01 

8.000E+00 1.823E-04 1.510E-03 7.403E-04 1.717E-01 7.968E-01 4.176E-01 

9.000E+00 2.730E-04 1.720E-03 8.702E-04 3.248E-01 1.848E+00 8.613E-01 

1.000E+01 3.911E-04 2.432E-03 1.186E-03 4.778E-01 2.900E+00 1.305E+00 

1.100E+01 5.092E-04 3.144E-03 1.502E-03 1.231E+00 5.957E+00 3.058E+00 

1.200E+01 1.138E-03 6.180E-03 3.008E-03 1.983E+00 9.015E+00 4.811E+00 

1.300E+01 1.767E-03 9.217E-03 4.514E-03 3.476E+00 1.335E+01 7.934E+00 

1.400E+01 3.628E-03 2.787E-02 1.115E-02 4.969E+00 1.768E+01 1.106E+01 

1.500E+01 5.489E-03 4.652E-02 1.778E-02 5.939E+00 2.086E+01 1.315E+01 

1.600E+01 1.779E-02 7.653E-02 4.228E-02 6.909E+00 2.404E+01 1.524E+01 

1.700E+01 3.009E-02 1.065E-01 6.678E-02 7.008E+00 2.438E+01 1.546E+01 

1.800E+01 3.608E-02 1.264E-01 7.978E-02 7.107E+00 2.472E+01 1.567E+01 

1.900E+01 4.207E-02 1.463E-01 9.277E-02 7.235E+00 2.516E+01 1.594E+01 

2.000E+01 4.248E-02 1.477E-01 9.367E-02 7.362E+00 2.559E+01 1.622E+01 

2.100E+01 4.289E-02 1.492E-01 9.456E-02 7.660E+00 2.662E+01 1.687E+01 

2.200E+01 4.314E-02 1.501E-01 9.512E-02 7.958E+00 2.765E+01 1.753E+01 

2.300E+01 4.339E-02 1.510E-01 9.567E-02 8.271E+00 2.876E+01 1.823E+01 

2.400E+01 4.410E-02 1.532E-01 9.714E-02 8.584E+00 2.987E+01 1.894E+01 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-24 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

NOE3 NOE4 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 4.706E-05 1.207E-03 5.453E-04 

-4.000E+00 1.701E-03 4.364E-02 1.971E-02 4.706E-05 1.207E-03 5.453E-04 

-3.000E+00 6.909E-03 1.772E-01 8.006E-02 4.706E-05 1.207E-03 5.453E-04 

-2.000E+00 1.212E-02 3.108E-01 1.404E-01 4.706E-05 1.207E-03 5.453E-04 

-1.000E+00 2.853E-02 4.161E-01 1.948E-01 4.706E-04 1.207E-02 5.453E-03 

0.000E+00 4.494E-02 5.213E-01 2.492E-01 5.929E-04 1.521E-02 6.870E-03 

1.000E+00 1.469E-01 9.374E-01 4.709E-01 7.153E-04 1.835E-02 8.288E-03 

2.000E+00 2.490E-01 1.353E+00 6.926E-01 1.205E-03 3.090E-02 1.396E-02 

3.000E+00 4.734E-01 2.427E+00 1.230E+00 1.694E-03 4.346E-02 1.963E-02 

4.000E+00 6.978E-01 3.501E+00 1.768E+00 2.542E-03 6.140E-02 2.496E-02 

5.000E+00 1.372E+00 5.405E+00 3.149E+00 3.389E-03 7.935E-02 3.029E-02 

6.000E+00 2.045E+00 7.308E+00 4.531E+00 3.340E-02 2.905E-01 1.184E-01 

7.000E+00 2.648E+00 9.288E+00 5.825E+00 6.340E-02 5.016E-01 2.066E-01 

8.000E+00 3.250E+00 1.127E+01 7.118E+00 2.399E-01 1.236E+00 6.146E-01 

9.000E+00 3.835E+00 1.367E+01 8.484E+00 4.163E-01 1.970E+00 1.023E+00 

1.000E+01 4.420E+00 1.606E+01 9.850E+00 7.026E-01 2.700E+00 1.599E+00 

1.100E+01 5.803E+00 2.320E+01 1.341E+01 9.889E-01 3.430E+00 2.176E+00 

1.200E+01 7.187E+00 3.034E+01 1.696E+01 1.017E+00 3.531E+00 2.240E+00 

1.300E+01 1.224E+01 4.814E+01 2.821E+01 1.046E+00 3.632E+00 2.304E+00 

1.400E+01 1.730E+01 6.594E+01 3.945E+01 1.078E+00 3.777E+00 2.383E+00 

1.500E+01 2.389E+01 8.593E+01 5.329E+01 1.111E+00 3.923E+00 2.463E+00 

1.600E+01 3.047E+01 1.059E+02 6.713E+01 1.285E+00 4.632E+00 2.870E+00 

1.700E+01 3.231E+01 1.124E+02 7.121E+01 1.459E+00 5.341E+00 3.278E+00 

1.800E+01 3.414E+01 1.188E+02 7.529E+01 1.919E+00 6.806E+00 4.257E+00 

1.900E+01 3.481E+01 1.211E+02 7.676E+01 2.380E+00 8.271E+00 5.236E+00 

2.000E+01 3.547E+01 1.234E+02 7.823E+01 2.581E+00 8.974E+00 5.686E+00 

2.100E+01 3.640E+01 1.265E+02 8.023E+01 2.783E+00 9.678E+00 6.137E+00 

2.200E+01 3.733E+01 1.297E+02 8.223E+01 2.808E+00 9.764E+00 6.191E+00 

2.300E+01 3.894E+01 1.354E+02 8.581E+01 2.833E+00 9.851E+00 6.245E+00 

2.400E+01 4.054E+01 1.410E+02 8.939E+01 2.866E+00 9.965E+00 6.317E+00 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-25 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

NOE5 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 5.058E-02 1.297E+00 5.861E-01 

-8.000E+00 7.701E-02 1.975E+00 8.923E-01 

-7.000E+00 1.034E-01 2.653E+00 1.199E+00 

-6.000E+00 1.383E-01 3.539E+00 1.599E+00 

-5.000E+00 1.731E-01 4.424E+00 1.998E+00 

-4.000E+00 2.592E-01 5.395E+00 2.466E+00 

-3.000E+00 3.453E-01 6.367E+00 2.934E+00 

-2.000E+00 8.712E-01 8.046E+00 3.905E+00 

-1.000E+00 1.397E+00 9.724E+00 4.875E+00 

0.000E+00 4.062E+00 1.869E+01 1.034E+01 

1.000E+00 6.726E+00 2.765E+01 1.580E+01 

2.000E+00 1.219E+01 4.612E+01 2.762E+01 

3.000E+00 1.765E+01 6.458E+01 3.944E+01 

4.000E+00 2.417E+01 8.573E+01 5.337E+01 

5.000E+00 3.069E+01 1.069E+02 6.729E+01 

6.000E+00 3.720E+01 1.318E+02 8.222E+01 

7.000E+00 4.370E+01 1.567E+02 9.715E+01 

8.000E+00 5.982E+01 2.285E+02 1.362E+02 

9.000E+00 7.593E+01 3.003E+02 1.752E+02 

1.000E+01 1.179E+02 4.373E+02 2.659E+02 

1.100E+01 1.598E+02 5.742E+02 3.566E+02 

1.200E+01 2.071E+02 7.314E+02 4.587E+02 

1.300E+01 2.544E+02 8.886E+02 5.608E+02 

1.400E+01 2.965E+02 1.033E+03 6.531E+02 

1.500E+01 3.387E+02 1.177E+03 7.453E+02 

1.600E+01 3.716E+02 1.291E+03 8.180E+02 

1.700E+01 4.046E+02 1.405E+03 8.907E+02 

1.800E+01 4.247E+02 1.475E+03 9.352E+02 

1.900E+01 4.447E+02 1.545E+03 9.797E+02 

2.000E+01 4.595E+02 1.597E+03 1.012E+03 

2.100E+01 4.742E+02 1.649E+03 1.045E+03 

2.200E+01 4.859E+02 1.689E+03 1.071E+03 

2.300E+01 4.976E+02 1.730E+03 1.097E+03 

2.400E+01 5.084E+02 1.768E+03 1.121E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-26 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

OM1 OM2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 4.014E-02 7.398E-01 3.407E-01 

-1.100E+01 1.648E-05 1.286E-03 3.144E-04 5.766E-02 9.349E-01 4.348E-01 

-1.000E+01 3.295E-05 2.572E-03 6.288E-04 2.091E-01 3.059E+00 1.091E+00 

-9.000E+00 2.424E-02 8.533E-02 5.353E-02 3.606E-01 5.184E+00 1.747E+00 

-8.000E+00 4.845E-02 1.681E-01 1.064E-01 2.156E+00 1.065E+01 5.460E+00 

-7.000E+00 6.112E-02 2.206E-01 1.360E-01 3.952E+00 1.611E+01 9.173E+00 

-6.000E+00 7.379E-02 2.732E-01 1.656E-01 6.342E+00 2.331E+01 1.418E+01 

-5.000E+00 1.209E-01 4.813E-01 2.835E-01 8.732E+00 3.051E+01 1.919E+01 

-4.000E+00 1.681E-01 6.895E-01 4.014E-01 1.053E+01 3.660E+01 2.310E+01 

-3.000E+00 2.134E-01 1.044E+00 5.740E-01 1.234E+01 4.270E+01 2.702E+01 

-2.000E+00 2.588E-01 1.398E+00 7.466E-01 1.424E+01 4.934E+01 3.122E+01 

-1.000E+00 2.588E-01 1.398E+00 7.466E-01 1.615E+01 5.599E+01 3.543E+01 

0.000E+00 2.588E-01 1.398E+00 7.466E-01 1.927E+01 7.023E+01 4.316E+01 

1.000E+00 1.071E+00 6.949E+00 3.228E+00 2.240E+01 8.447E+01 5.090E+01 

2.000E+00 2.796E+00 1.382E+01 7.100E+00 3.241E+01 1.171E+02 7.242E+01 

3.000E+00 4.520E+00 2.070E+01 1.097E+01 4.243E+01 1.496E+02 9.395E+01 

4.000E+00 7.901E+00 3.010E+01 1.788E+01 5.211E+01 1.824E+02 1.150E+02 

5.000E+00 1.128E+01 3.951E+01 2.479E+01 6.180E+01 2.152E+02 1.361E+02 

6.000E+00 1.394E+01 4.866E+01 3.060E+01 7.093E+01 2.466E+02 1.561E+02 

7.000E+00 1.659E+01 5.781E+01 3.641E+01 8.005E+01 2.779E+02 1.762E+02 

8.000E+00 2.000E+01 6.983E+01 4.396E+01 8.631E+01 2.997E+02 1.900E+02 

9.000E+00 2.341E+01 8.185E+01 5.150E+01 9.256E+01 3.214E+02 2.038E+02 

1.000E+01 2.875E+01 1.019E+02 6.367E+01 9.722E+01 3.375E+02 2.140E+02 

1.100E+01 3.408E+01 1.219E+02 7.584E+01 1.019E+02 3.536E+02 2.242E+02 

1.200E+01 4.513E+01 1.618E+02 1.006E+02 1.082E+02 3.758E+02 2.383E+02 

1.300E+01 5.618E+01 2.018E+02 1.254E+02 1.146E+02 3.980E+02 2.523E+02 

1.400E+01 7.389E+01 2.626E+02 1.642E+02 1.221E+02 4.242E+02 2.689E+02 

1.500E+01 9.161E+01 3.234E+02 2.030E+02 1.296E+02 4.505E+02 2.855E+02 

1.600E+01 1.127E+02 3.952E+02 2.490E+02 1.338E+02 4.654E+02 2.950E+02 

1.700E+01 1.337E+02 4.670E+02 2.950E+02 1.380E+02 4.803E+02 3.044E+02 

1.800E+01 1.536E+02 5.347E+02 3.384E+02 1.400E+02 4.872E+02 3.089E+02 

1.900E+01 1.734E+02 6.025E+02 3.818E+02 1.420E+02 4.942E+02 3.133E+02 

2.000E+01 1.867E+02 6.483E+02 4.109E+02 1.437E+02 4.999E+02 3.169E+02 

2.100E+01 1.999E+02 6.942E+02 4.401E+02 1.453E+02 5.055E+02 3.205E+02 

2.200E+01 2.093E+02 7.272E+02 4.609E+02 1.472E+02 5.122E+02 3.247E+02 

2.300E+01 2.187E+02 7.601E+02 4.818E+02 1.491E+02 5.189E+02 3.290E+02 

2.400E+01 2.261E+02 7.858E+02 4.981E+02 1.507E+02 5.246E+02 3.325E+02 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-27 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

OM3 OM4 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 3.576E-04 9.174E-03 4.144E-03 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 7.153E-04 1.835E-02 8.288E-03 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 4.071E-03 1.126E-01 3.988E-02 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 7.426E-03 2.068E-01 7.147E-02 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 6.749E-02 5.597E-01 2.318E-01 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 1.276E-01 9.127E-01 3.921E-01 

-5.000E+00 4.656E-03 9.461E-02 3.208E-02 3.614E-01 1.490E+00 8.471E-01 

-4.000E+00 9.312E-03 1.892E-01 6.416E-02 5.952E-01 2.067E+00 1.302E+00 

-3.000E+00 1.052E-01 6.077E-01 3.092E-01 7.235E-01 2.533E+00 1.589E+00 

-2.000E+00 2.010E-01 1.026E+00 5.542E-01 8.517E-01 3.000E+00 1.876E+00 

-1.000E+00 3.333E-01 2.665E+00 1.322E+00 1.049E+00 3.983E+00 2.373E+00 

0.000E+00 4.656E-01 4.304E+00 2.090E+00 1.246E+00 4.967E+00 2.869E+00 

1.000E+00 7.944E-01 8.308E+00 3.697E+00 1.874E+00 8.033E+00 4.430E+00 

2.000E+00 1.123E+00 1.231E+01 5.305E+00 2.502E+00 1.110E+01 5.992E+00 

3.000E+00 4.897E+00 2.978E+01 1.405E+01 4.446E+00 1.745E+01 1.018E+01 

4.000E+00 8.670E+00 4.725E+01 2.279E+01 6.390E+00 2.381E+01 1.436E+01 

5.000E+00 1.896E+01 7.549E+01 4.359E+01 9.482E+00 3.406E+01 2.106E+01 

6.000E+00 2.925E+01 1.037E+02 6.439E+01 1.257E+01 4.431E+01 2.775E+01 

7.000E+00 3.680E+01 1.294E+02 8.087E+01 1.579E+01 5.526E+01 3.476E+01 

8.000E+00 4.435E+01 1.550E+02 9.735E+01 1.901E+01 6.621E+01 4.177E+01 

9.000E+00 5.438E+01 1.901E+02 1.194E+02 2.298E+01 8.057E+01 5.066E+01 

1.000E+01 6.441E+01 2.253E+02 1.415E+02 2.695E+01 9.493E+01 5.955E+01 

1.100E+01 8.052E+01 2.907E+02 1.794E+02 3.383E+01 1.197E+02 7.492E+01 

1.200E+01 9.664E+01 3.561E+02 2.174E+02 4.071E+01 1.445E+02 9.030E+01 

1.300E+01 1.369E+02 4.955E+02 3.059E+02 5.140E+01 1.804E+02 1.135E+02 

1.400E+01 1.771E+02 6.349E+02 3.945E+02 6.209E+01 2.163E+02 1.366E+02 

1.500E+01 2.322E+02 8.198E+02 5.141E+02 7.253E+01 2.524E+02 1.596E+02 

1.600E+01 2.873E+02 1.005E+03 6.338E+02 8.296E+01 2.886E+02 1.825E+02 

1.700E+01 3.419E+02 1.192E+03 7.529E+02 9.452E+01 3.286E+02 2.080E+02 

1.800E+01 3.966E+02 1.378E+03 8.721E+02 1.061E+02 3.686E+02 2.334E+02 

1.900E+01 4.554E+02 1.585E+03 1.002E+03 1.170E+02 4.065E+02 2.574E+02 

2.000E+01 5.143E+02 1.791E+03 1.132E+03 1.279E+02 4.444E+02 2.815E+02 

2.100E+01 5.927E+02 2.062E+03 1.305E+03 1.394E+02 4.842E+02 3.068E+02 

2.200E+01 6.711E+02 2.333E+03 1.477E+03 1.509E+02 5.240E+02 3.322E+02 

2.300E+01 7.454E+02 2.590E+03 1.640E+03 1.585E+02 5.504E+02 3.489E+02 

2.400E+01 8.197E+02 2.846E+03 1.804E+03 1.660E+02 5.768E+02 3.657E+02 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-28 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

OM5 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 4.621E-02 2.368E-01 1.281E-01 

-3.000E+00 9.242E-02 4.735E-01 2.561E-01 

-2.000E+00 3.024E-01 3.900E+00 1.837E+00 

-1.000E+00 5.124E-01 7.326E+00 3.418E+00 

0.000E+00 8.923E-01 1.230E+01 5.751E+00 

1.000E+00 1.272E+00 1.727E+01 8.085E+00 

2.000E+00 2.051E+00 2.555E+01 1.210E+01 

3.000E+00 2.830E+00 3.384E+01 1.611E+01 

4.000E+00 7.114E+00 6.206E+01 2.813E+01 

5.000E+00 1.140E+01 9.028E+01 4.014E+01 

6.000E+00 3.477E+01 1.669E+02 8.835E+01 

7.000E+00 5.814E+01 2.435E+02 1.365E+02 

8.000E+00 9.600E+01 3.695E+02 2.184E+02 

9.000E+00 1.339E+02 4.956E+02 3.002E+02 

1.000E+01 1.875E+02 6.704E+02 4.153E+02 

1.100E+01 2.412E+02 8.452E+02 5.303E+02 

1.200E+01 2.983E+02 1.046E+03 6.565E+02 

1.300E+01 3.554E+02 1.246E+03 7.828E+02 

1.400E+01 4.512E+02 1.646E+03 1.011E+03 

1.500E+01 5.470E+02 2.047E+03 1.239E+03 

1.600E+01 7.893E+02 2.861E+03 1.767E+03 

1.700E+01 1.032E+03 3.676E+03 2.295E+03 

1.800E+01 1.323E+03 4.669E+03 2.931E+03 

1.900E+01 1.615E+03 5.661E+03 3.568E+03 

2.000E+01 1.928E+03 6.729E+03 4.252E+03 

2.100E+01 2.241E+03 7.797E+03 4.935E+03 

2.200E+01 2.488E+03 8.650E+03 5.478E+03 

2.300E+01 2.736E+03 9.503E+03 6.021E+03 

2.400E+01 2.972E+03 1.032E+04 6.543E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-29 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

SB1 SB3 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

1.000E+00 1.538E-01 2.010E+00 9.435E-01 3.095E-02 4.228E-01 1.978E-01 

2.000E+00 3.075E-01 4.020E+00 1.887E+00 6.190E-02 8.456E-01 3.957E-01 

3.000E+00 3.249E-01 4.141E+00 1.948E+00 1.954E-01 2.552E+00 1.199E+00 

4.000E+00 3.422E-01 4.262E+00 2.008E+00 3.289E-01 4.259E+00 2.001E+00 

5.000E+00 6.525E-01 1.331E+01 4.361E+00 3.924E-01 4.376E+00 2.085E+00 

6.000E+00 9.628E-01 2.236E+01 6.713E+00 4.559E-01 4.493E+00 2.169E+00 

7.000E+00 9.138E+00 4.768E+01 2.373E+01 3.702E+00 1.535E+01 9.133E+00 

8.000E+00 1.731E+01 7.300E+01 4.075E+01 6.948E+00 2.620E+01 1.610E+01 

9.000E+00 2.770E+01 1.029E+02 6.238E+01 1.953E+01 6.875E+01 4.335E+01 

1.000E+01 3.809E+01 1.329E+02 8.401E+01 3.212E+01 1.113E+02 7.061E+01 

1.100E+01 3.864E+01 1.344E+02 8.514E+01 3.212E+01 1.113E+02 7.061E+01 

1.200E+01 3.918E+01 1.360E+02 8.627E+01 3.212E+01 1.113E+02 7.061E+01 

1.300E+01 3.918E+01 1.360E+02 8.628E+01 3.297E+01 1.285E+02 7.539E+01 

1.400E+01 3.918E+01 1.360E+02 8.628E+01 3.382E+01 1.457E+02 8.017E+01 

1.500E+01 4.295E+01 1.905E+02 1.031E+02 5.103E+01 2.482E+02 1.268E+02 

1.600E+01 4.671E+01 2.449E+02 1.199E+02 6.824E+01 3.508E+02 1.734E+02 

1.700E+01 9.819E+01 3.885E+02 2.261E+02 1.252E+02 4.919E+02 2.873E+02 

1.800E+01 1.497E+02 5.322E+02 3.324E+02 1.822E+02 6.330E+02 4.012E+02 

1.900E+01 1.809E+02 6.359E+02 4.004E+02 1.890E+02 6.577E+02 4.168E+02 

2.000E+01 2.122E+02 7.397E+02 4.685E+02 1.959E+02 6.824E+02 4.323E+02 

2.100E+01 2.137E+02 7.448E+02 4.718E+02 1.959E+02 6.824E+02 4.323E+02 

2.200E+01 2.152E+02 7.499E+02 4.751E+02 1.959E+02 6.824E+02 4.323E+02 

2.300E+01 2.165E+02 7.535E+02 4.774E+02 1.967E+02 6.852E+02 4.341E+02 

2.400E+01 2.178E+02 7.572E+02 4.798E+02 1.976E+02 6.880E+02 4.358E+02 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-30 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

SB4 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 

1.000E+00 4.429E-03 9.593E-02 4.364E-02 

2.000E+00 8.859E-03 1.919E-01 8.728E-02 

3.000E+00 5.590E-02 1.075E+00 4.924E-01 

4.000E+00 1.029E-01 1.958E+00 8.975E-01 

5.000E+00 2.036E-01 2.272E+00 1.079E+00 

6.000E+00 3.043E-01 2.586E+00 1.261E+00 

7.000E+00 7.299E-01 4.183E+00 2.123E+00 

8.000E+00 1.156E+00 5.779E+00 2.984E+00 

9.000E+00 2.260E+00 9.830E+00 5.411E+00 

1.000E+01 3.365E+00 1.388E+01 7.838E+00 

1.100E+01 5.317E+00 1.970E+01 1.192E+01 

1.200E+01 7.269E+00 2.552E+01 1.601E+01 

1.300E+01 8.792E+00 3.376E+01 1.995E+01 

1.400E+01 1.031E+01 4.200E+01 2.389E+01 

1.500E+01 1.113E+01 7.379E+01 3.162E+01 

1.600E+01 1.194E+01 1.056E+02 3.936E+01 

1.700E+01 1.478E+01 1.122E+02 4.495E+01 

1.800E+01 1.761E+01 1.189E+02 5.055E+01 

1.900E+01 4.328E+01 1.793E+02 1.012E+02 

2.000E+01 6.895E+01 2.397E+02 1.519E+02 

2.100E+01 7.069E+01 2.459E+02 1.558E+02 

2.200E+01 7.242E+01 2.522E+02 1.598E+02 

2.300E+01 7.261E+01 2.528E+02 1.602E+02 

2.400E+01 7.280E+01 2.535E+02 1.606E+02 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-31 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

JE1 JE2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 5.956E-02 1.528E+00 6.901E-01 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 6.700E-02 1.719E+00 7.764E-01 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 7.445E-02 1.910E+00 8.626E-01 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 5.770E-01 6.503E+00 2.494E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 1.080E+00 1.110E+01 4.126E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 5.812E+00 2.505E+01 1.387E+01 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 1.055E+01 3.899E+01 2.362E+01 

-4.000E+00 1.334E-01 5.056E-01 3.016E-01 1.669E+01 6.096E+01 3.732E+01 

-3.000E+00 2.668E-01 1.011E+00 6.031E-01 2.284E+01 8.292E+01 5.103E+01 

-2.000E+00 7.894E-01 2.776E+00 1.737E+00 2.842E+01 1.068E+02 6.460E+01 

-1.000E+00 1.312E+00 4.540E+00 2.870E+00 3.400E+01 1.306E+02 7.816E+01 

0.000E+00 1.820E+00 6.716E+00 4.109E+00 4.417E+01 2.372E+02 1.153E+02 

1.000E+00 2.329E+00 8.892E+00 5.347E+00 5.435E+01 3.438E+02 1.525E+02 

2.000E+00 3.048E+00 1.215E+01 7.170E+00 1.463E+02 6.670E+02 3.552E+02 

3.000E+00 3.768E+00 1.541E+01 8.993E+00 2.382E+02 9.903E+02 5.580E+02 

4.000E+00 5.941E+00 3.590E+01 1.696E+01 3.787E+02 1.412E+03 8.521E+02 

5.000E+00 8.114E+00 5.639E+01 2.493E+01 5.191E+02 1.834E+03 1.146E+03 

6.000E+00 2.066E+01 9.780E+01 5.153E+01 6.277E+02 2.191E+03 1.381E+03 

7.000E+00 3.320E+01 1.392E+02 7.813E+01 7.363E+02 2.548E+03 1.616E+03 

8.000E+00 5.805E+01 2.252E+02 1.324E+02 8.013E+02 2.774E+03 1.758E+03 

9.000E+00 8.290E+01 3.111E+02 1.866E+02 8.663E+02 3.000E+03 1.900E+03 

1.000E+01 1.204E+02 4.358E+02 2.679E+02 1.009E+03 3.710E+03 2.266E+03 

1.100E+01 1.578E+02 5.606E+02 3.492E+02 1.153E+03 4.420E+03 2.632E+03 

1.200E+01 2.035E+02 7.207E+02 4.499E+02 1.689E+03 6.157E+03 3.785E+03 

1.300E+01 2.492E+02 8.808E+02 5.506E+02 2.225E+03 7.894E+03 4.937E+03 

1.400E+01 3.216E+02 1.148E+03 7.141E+02 2.713E+03 9.515E+03 5.995E+03 

1.500E+01 3.941E+02 1.416E+03 8.775E+02 3.202E+03 1.114E+04 7.053E+03 

1.600E+01 5.299E+02 1.895E+03 1.179E+03 3.479E+03 1.210E+04 7.663E+03 

1.700E+01 6.656E+02 2.375E+03 1.480E+03 3.756E+03 1.306E+04 8.274E+03 

1.800E+01 8.657E+02 3.066E+03 1.920E+03 3.914E+03 1.360E+04 8.622E+03 

1.900E+01 1.066E+03 3.757E+03 2.360E+03 4.073E+03 1.415E+04 8.971E+03 

2.000E+01 1.308E+03 4.581E+03 2.888E+03 4.238E+03 1.472E+04 9.333E+03 

2.100E+01 1.551E+03 5.404E+03 3.416E+03 4.403E+03 1.529E+04 9.696E+03 

2.200E+01 1.791E+03 6.234E+03 3.944E+03 4.581E+03 1.592E+04 1.009E+04 

2.300E+01 2.031E+03 7.064E+03 4.472E+03 4.760E+03 1.655E+04 1.049E+04 

2.400E+01 2.249E+03 7.813E+03 4.950E+03 4.861E+03 1.691E+04 1.072E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-32 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

JE3 

Elevation (ft) 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 4.116E-02 7.489E-01 3.454E-01 

-8.000E+00 8.231E-02 1.498E+00 6.907E-01 

-7.000E+00 1.362E+00 1.153E+01 4.992E+00 

-6.000E+00 2.641E+00 2.155E+01 9.293E+00 

-5.000E+00 1.268E+01 5.775E+01 3.178E+01 

-4.000E+00 2.272E+01 9.395E+01 5.426E+01 

-3.000E+00 3.748E+01 1.668E+02 9.411E+01 

-2.000E+00 5.224E+01 2.397E+02 1.340E+02 

-1.000E+00 6.749E+01 3.150E+02 1.720E+02 

0.000E+00 8.274E+01 3.903E+02 2.100E+02 

1.000E+00 1.594E+02 8.534E+02 4.154E+02 

2.000E+00 2.362E+02 1.316E+03 6.207E+02 

3.000E+00 4.798E+02 1.997E+03 1.124E+03 

4.000E+00 7.235E+02 2.678E+03 1.627E+03 

5.000E+00 1.018E+03 3.605E+03 2.250E+03 

6.000E+00 1.312E+03 4.532E+03 2.873E+03 

7.000E+00 1.405E+03 4.863E+03 3.081E+03 

8.000E+00 1.499E+03 5.193E+03 3.289E+03 

9.000E+00 1.746E+03 6.515E+03 3.942E+03 

1.000E+01 1.992E+03 7.837E+03 4.595E+03 

1.100E+01 3.152E+03 1.208E+04 7.212E+03 

1.200E+01 4.312E+03 1.632E+04 9.829E+03 

1.300E+01 6.108E+03 2.189E+04 1.362E+04 

1.400E+01 7.904E+03 2.745E+04 1.740E+04 

1.500E+01 8.391E+03 2.917E+04 1.849E+04 

1.600E+01 8.878E+03 3.088E+04 1.958E+04 

1.700E+01 9.071E+03 3.155E+04 2.000E+04 

1.800E+01 9.264E+03 3.222E+04 2.043E+04 

1.900E+01 9.508E+03 3.305E+04 2.095E+04 

2.000E+01 9.752E+03 3.387E+04 2.147E+04 

2.100E+01 1.024E+04 3.559E+04 2.256E+04 

2.200E+01 1.073E+04 3.730E+04 2.364E+04 

2.300E+01 1.102E+04 3.832E+04 2.429E+04 

2.400E+01 1.130E+04 3.933E+04 2.493E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-33 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

JW1 JW2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 8.278E-03 1.981E-01 8.970E-02 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 1.656E-02 3.962E-01 1.794E-01 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 2.174E-02 4.140E-01 1.896E-01 3.882E-05 9.958E-04 4.499E-04 

-5.000E+00 2.692E-02 4.318E-01 1.999E-01 7.765E-05 1.992E-03 8.997E-04 

-4.000E+00 9.405E-02 1.512E+00 6.996E-01 2.114E-03 5.033E-02 2.280E-02 

-3.000E+00 1.612E-01 2.592E+00 1.199E+00 4.150E-03 9.866E-02 4.471E-02 

-2.000E+00 2.814E-01 3.096E+00 1.478E+00 3.605E-02 7.686E-01 3.500E-01 

-1.000E+00 4.017E-01 3.599E+00 1.757E+00 6.796E-02 1.438E+00 6.554E-01 

0.000E+00 1.156E+00 8.771E+00 3.791E+00 2.180E-01 3.649E+00 1.684E+00 

1.000E+00 1.911E+00 1.394E+01 5.824E+00 3.681E-01 5.860E+00 2.713E+00 

2.000E+00 5.236E+00 2.462E+01 1.288E+01 6.348E-01 7.903E+00 3.723E+00 

3.000E+00 8.561E+00 3.531E+01 1.994E+01 9.015E-01 9.946E+00 4.733E+00 

4.000E+00 1.419E+01 5.197E+01 3.167E+01 1.537E+00 1.318E+01 6.328E+00 

5.000E+00 1.983E+01 6.863E+01 4.341E+01 2.172E+00 1.641E+01 7.923E+00 

6.000E+00 2.162E+01 7.487E+01 4.732E+01 5.744E+00 3.265E+01 1.596E+01 

7.000E+00 2.342E+01 8.112E+01 5.123E+01 9.316E+00 4.888E+01 2.400E+01 

8.000E+00 2.808E+01 1.058E+02 6.330E+01 1.755E+01 7.099E+01 4.066E+01 

9.000E+00 3.275E+01 1.304E+02 7.537E+01 2.578E+01 9.310E+01 5.732E+01 

1.000E+01 5.405E+01 2.185E+02 1.252E+02 3.399E+01 1.192E+02 7.480E+01 

1.100E+01 7.536E+01 3.065E+02 1.751E+02 4.220E+01 1.453E+02 9.228E+01 

1.200E+01 1.185E+02 4.340E+02 2.650E+02 4.580E+01 1.614E+02 1.009E+02 

1.300E+01 1.616E+02 5.614E+02 3.549E+02 4.940E+01 1.775E+02 1.095E+02 

1.400E+01 1.768E+02 6.127E+02 3.881E+02 6.504E+01 2.909E+02 1.564E+02 

1.500E+01 1.919E+02 6.639E+02 4.213E+02 8.068E+01 4.043E+02 2.032E+02 

1.600E+01 2.041E+02 7.073E+02 4.481E+02 1.661E+02 6.714E+02 3.865E+02 

1.700E+01 2.163E+02 7.507E+02 4.749E+02 2.515E+02 9.386E+02 5.698E+02 

1.800E+01 2.537E+02 8.933E+02 5.610E+02 3.440E+02 1.227E+03 7.652E+02 

1.900E+01 2.911E+02 1.036E+03 6.471E+02 4.365E+02 1.515E+03 9.606E+02 

2.000E+01 3.706E+02 1.302E+03 8.188E+02 4.569E+02 1.588E+03 1.007E+03 

2.100E+01 4.500E+02 1.568E+03 9.905E+02 4.773E+02 1.660E+03 1.053E+03 

2.200E+01 5.161E+02 1.794E+03 1.136E+03 4.828E+02 1.680E+03 1.065E+03 

2.300E+01 5.822E+02 2.021E+03 1.281E+03 4.883E+02 1.699E+03 1.077E+03 

2.400E+01 6.192E+02 2.150E+03 1.363E+03 5.016E+02 1.743E+03 1.105E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-34 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

JW3 JW4 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 5.118E-04 8.286E-03 3.860E-03 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 1.024E-03 1.657E-02 7.719E-03 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 1.270E-02 1.662E-01 7.856E-02 

-7.000E+00 3.859E-04 9.898E-03 4.471E-03 2.438E-02 3.159E-01 1.494E-01 

-6.000E+00 7.717E-04 1.980E-02 8.942E-03 1.162E-01 1.247E+00 5.992E-01 

-5.000E+00 9.314E-03 2.051E-01 8.210E-02 2.080E-01 2.179E+00 1.049E+00 

-4.000E+00 1.786E-02 3.904E-01 1.553E-01 6.595E-01 5.946E+00 2.909E+00 

-3.000E+00 2.028E-01 1.843E+00 7.957E-01 1.111E+00 9.714E+00 4.769E+00 

-2.000E+00 3.878E-01 3.295E+00 1.436E+00 3.270E+00 2.504E+01 1.251E+01 

-1.000E+00 1.882E+00 9.998E+00 5.275E+00 5.428E+00 4.037E+01 2.025E+01 

0.000E+00 3.376E+00 1.670E+01 9.114E+00 9.568E+00 6.253E+01 3.207E+01 

1.000E+00 5.791E+00 3.119E+01 1.665E+01 1.371E+01 8.469E+01 4.388E+01 

2.000E+00 8.207E+00 4.568E+01 2.418E+01 2.268E+01 1.203E+02 6.404E+01 

3.000E+00 1.107E+01 6.386E+01 3.350E+01 3.165E+01 1.559E+02 8.420E+01 

4.000E+00 1.393E+01 8.204E+01 4.282E+01 5.403E+01 2.448E+02 1.340E+02 

5.000E+00 2.108E+01 1.193E+02 6.217E+01 7.640E+01 3.337E+02 1.837E+02 

6.000E+00 2.823E+01 1.566E+02 8.153E+01 1.309E+02 5.288E+02 3.035E+02 

7.000E+00 5.501E+01 2.599E+02 1.398E+02 1.854E+02 7.239E+02 4.233E+02 

8.000E+00 8.180E+01 3.632E+02 1.980E+02 2.764E+02 1.001E+03 6.143E+02 

9.000E+00 1.402E+02 5.479E+02 3.213E+02 3.673E+02 1.279E+03 8.052E+02 

1.000E+01 1.987E+02 7.326E+02 4.447E+02 4.455E+02 1.566E+03 9.815E+02 

1.100E+01 2.749E+02 9.885E+02 6.104E+02 5.237E+02 1.854E+03 1.158E+03 

1.200E+01 3.511E+02 1.244E+03 7.761E+02 6.807E+02 2.458E+03 1.518E+03 

1.300E+01 4.560E+02 1.627E+03 1.012E+03 8.378E+02 3.062E+03 1.877E+03 

1.400E+01 5.610E+02 2.010E+03 1.248E+03 1.179E+03 4.524E+03 2.691E+03 

1.500E+01 7.497E+02 2.686E+03 1.668E+03 1.521E+03 5.986E+03 3.505E+03 

1.600E+01 9.384E+02 3.362E+03 2.088E+03 2.289E+03 8.342E+03 5.127E+03 

1.700E+01 1.257E+03 4.606E+03 2.824E+03 3.057E+03 1.070E+04 6.748E+03 

1.800E+01 1.575E+03 5.851E+03 3.560E+03 3.608E+03 1.257E+04 7.948E+03 

1.900E+01 2.211E+03 7.935E+03 4.930E+03 4.158E+03 1.445E+04 9.149E+03 

2.000E+01 2.846E+03 1.002E+04 6.299E+03 4.592E+03 1.595E+04 1.010E+04 

2.100E+01 3.445E+03 1.204E+04 7.601E+03 5.025E+03 1.744E+04 1.106E+04 

2.200E+01 4.044E+03 1.407E+04 8.902E+03 5.301E+03 1.840E+04 1.167E+04 

2.300E+01 4.517E+03 1.570E+04 9.943E+03 5.577E+03 1.936E+04 1.227E+04 

2.400E+01 4.990E+03 1.733E+04 1.098E+04 5.872E+03 2.039E+04 1.293E+04 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-35 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Continued) 

PL11 

Elevation (ft) 5% 95% Mean 

-1.200E+01 2.828E-01 1.235E+00 7.010E-01 

-1.100E+01 3.505E-01 1.808E+00 9.773E-01 

-1.000E+01 4.181E-01 2.382E+00 1.254E+00 

-9.000E+00 4.645E-01 2.584E+00 1.368E+00 

-8.000E+00 5.109E-01 2.786E+00 1.483E+00 

-7.000E+00 7.870E-01 6.230E+00 2.589E+00 

-6.000E+00 1.063E+00 9.673E+00 3.695E+00 

-5.000E+00 4.147E+00 1.836E+01 9.993E+00 

-4.000E+00 7.231E+00 2.705E+01 1.629E+01 

-3.000E+00 1.020E+01 3.621E+01 2.254E+01 

-2.000E+00 1.316E+01 4.536E+01 2.878E+01 

-1.000E+00 1.438E+01 4.950E+01 3.143E+01 

0.000E+00 1.559E+01 5.364E+01 3.408E+01 

1.000E+00 1.720E+01 6.055E+01 3.794E+01 

2.000E+00 1.880E+01 6.746E+01 4.180E+01 

3.000E+00 2.640E+01 1.082E+02 6.146E+01 

4.000E+00 3.401E+01 1.489E+02 8.112E+01 

5.000E+00 5.866E+01 2.188E+02 1.320E+02 

6.000E+00 8.332E+01 2.887E+02 1.828E+02 

7.000E+00 9.352E+01 3.238E+02 2.052E+02 

8.000E+00 1.037E+02 3.590E+02 2.275E+02 

9.000E+00 1.147E+02 3.972E+02 2.516E+02 

1.000E+01 1.257E+02 4.355E+02 2.758E+02 

1.100E+01 1.408E+02 4.881E+02 3.090E+02 

1.200E+01 1.560E+02 5.408E+02 3.423E+02 

1.300E+01 1.774E+02 6.165E+02 3.898E+02 

1.400E+01 1.988E+02 6.922E+02 4.373E+02 

1.500E+01 2.290E+02 7.957E+02 5.038E+02 

1.600E+01 2.592E+02 8.992E+02 5.702E+02 

1.700E+01 2.754E+02 9.558E+02 6.057E+02 

1.800E+01 2.916E+02 1.012E+03 6.412E+02 

1.900E+01 3.250E+02 1.132E+03 7.157E+02 

2.000E+01 3.583E+02 1.252E+03 7.902E+02 

2.100E+01 4.152E+02 1.446E+03 9.149E+02 

2.200E+01 4.721E+02 1.640E+03 1.040E+03 

2.300E+01 5.036E+02 1.749E+03 1.109E+03 

2.400E+01 5.351E+02 1.859E+03 1.179E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-36 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-4 Current HPS – Stage-Fatality (Concluded) 

SC1 SC2 

Elevation (ft) 5% 95% Mean 5% 95% Mean 

-1.200E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.100E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+01 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-9.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-8.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-7.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-6.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-5.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-4.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-3.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-2.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

-1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

1.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 

2.000E+00 0.000E+00 0.000E+00 0.000E+00 3.718E-03 9.536E-02 4.308E-02 

3.000E+00 0.000E+00 0.000E+00 0.000E+00 7.435E-03 1.907E-01 8.615E-02 

4.000E+00 0.000E+00 0.000E+00 0.000E+00 6.044E-02 1.541E+00 6.961E-01 

5.000E+00 0.000E+00 0.000E+00 0.000E+00 1.135E-01 2.891E+00 1.306E+00 

6.000E+00 1.059E-05 2.716E-04 1.227E-04 4.328E-01 4.675E+00 2.238E+00 

7.000E+00 2.118E-05 5.432E-04 2.454E-04 7.521E-01 6.460E+00 3.169E+00 

8.000E+00 1.589E-02 5.967E-02 3.676E-02 1.890E+00 1.507E+01 6.764E+00 

9.000E+00 3.177E-02 1.188E-01 7.327E-02 3.027E+00 2.368E+01 1.036E+01 

1.000E+01 3.837E-01 1.330E+00 8.410E-01 9.488E+00 4.392E+01 2.362E+01 

1.100E+01 7.356E-01 2.541E+00 1.609E+00 1.595E+01 6.416E+01 3.689E+01 

1.200E+01 1.772E+00 6.096E+00 3.871E+00 2.669E+01 1.019E+02 6.047E+01 

1.300E+01 2.809E+00 9.652E+00 6.133E+00 3.742E+01 1.396E+02 8.405E+01 

1.400E+01 2.961E+00 1.018E+01 6.465E+00 5.407E+01 1.925E+02 1.194E+02 

1.500E+01 3.113E+00 1.071E+01 6.798E+00 7.071E+01 2.454E+02 1.548E+02 

1.600E+01 3.323E+00 1.243E+01 7.466E+00 8.778E+01 3.094E+02 1.936E+02 

1.700E+01 3.533E+00 1.416E+01 8.133E+00 1.048E+02 3.734E+02 2.325E+02 

1.800E+01 5.476E+00 3.828E+01 1.598E+01 1.486E+02 5.749E+02 3.396E+02 

1.900E+01 7.419E+00 6.240E+01 2.384E+01 1.924E+02 7.764E+02 4.468E+02 

2.000E+01 2.305E+01 9.842E+01 5.453E+01 3.065E+02 1.131E+03 6.896E+02 

2.100E+01 3.867E+01 1.345E+02 8.522E+01 4.206E+02 1.486E+03 9.324E+02 

2.200E+01 3.963E+01 1.378E+02 8.737E+01 5.287E+02 1.866E+03 1.172E+03 

2.300E+01 4.058E+01 1.412E+02 8.951E+01 6.369E+02 2.246E+03 1.411E+03 

2.400E+01 4.090E+01 1.423E+02 9.020E+01 7.825E+02 2.735E+03 1.726E+03 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-37 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

JE2 -12 0.0 0.0 

JE2 -11 0.0 0.0 

JE2 -10 0.0 0.0 

JE2 -9 0.9 0.9 

JE2 -8 1.0 1.0 

JE2 -7 2.4 2.4 

JE2 -6 8.1 8.1 

JE2 -5 52.0 52.0 

JE2 -4 470.7 470.7 

JE2 -3 2,190.3 2,190.3 

JE2 -2 3,394.0 3,394.0 

JE2 -1 3,857.5 3,857.5 

JE2 0 4,228.2 4,228.2 

JE2 1 4,506.3 4,504.7 

JE2 2 4,752.0 4,747.6 

JE2 3 4,994.2 4,985.7 

JE2 4 5,237.5 5,224.7 

JE2 5 5,499.4 5,485.1 

JE2 6 5,736.9 5,722.4 

JE2 7 5,864.8 5,850.0 

JE2 8 5,943.8 5,928.9 

JE2 9 6,079.7 6,064.5 

JE2 10 6,175.8 6,160.3 

JE2 11 6,215.7 6,199.9 

JE2 12 6,243.1 6,227.0 

JE2 13 6,262.7 6,246.4 

JE2 14 6,277.7 6,261.2 

JE2 15 6,286.6 6,269.9 

JE2 16 6,292.0 6,275.2 

JE2 17 6,295.9 6,279.0 

JE2 18 6,298.8 6,281.9 

JE2 19 6,300.2 6,283.3 

JE2 20 6,300.6 6,283.7 

JE2 21 6,300.7 6,283.8 

JE2 22 6,300.7 6,283.9 

JE2 23 6,300.7 6,283.9 

JE2 24 6,300.7 6,283.9 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-38 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

NOE1 -3 0.0 0.0 

NOE1 -2 0.0 0.0 

NOE1 -1 0.0 0.0 

NOE1 0 0.1 0.0 

NOE1 1 6.0 0.1 

NOE1 2 8.0 0.1 

NOE1 3 9.1 0.1 

NOE1 4 10.5 0.1 

NOE1 5 10.9 0.1 

NOE1 6 10.9 0.1 

NOE1 7 10.9 0.1 

NOE1 8 11.3 0.1 

NOE1 9 11.7 0.1 

NOE1 10 12.0 0.1 

NOE1 11 12.1 0.1 

NOE1 12 12.1 0.1 

NOE1 13 12.3 0.1 

NOE1 14 12.4 0.1 

NOE1 15 12.4 0.1 

NOE1 16 12.4 0.1 

NOE1 17 12.4 0.1 

NOE1 18 12.4 0.1 

NOE1 19 12.4 0.1 

NOE1 20 12.4 0.1 

NOE1 21 12.4 0.1 

NOE1 22 12.4 0.1 

NOE1 23 12.4 0.1 

NOE1 24 12.4 0.1 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-39 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

NOE2 -8 0.0 0.0 

NOE2 -7 0.0 0.0 

NOE2 -6 0.0 0.0 

NOE2 -5 0.5 0.0 

NOE2 -4 19.3 0.0 

NOE2 -3 29.0 0.0 

NOE2 -2 98.3 0.7 

NOE2 -1 114.5 0.9 

NOE2 0 116.3 0.9 

NOE2 1 120.6 1.0 

NOE2 2 123.3 1.0 

NOE2 3 124.1 1.1 

NOE2 4 126.1 1.1 

NOE2 5 127.6 1.1 

NOE2 6 134.5 1.1 

NOE2 7 138.6 1.2 

NOE2 8 139.8 1.2 

NOE2 9 140.8 1.2 

NOE2 10 142.0 1.2 

NOE2 11 142.5 1.2 

NOE2 12 142.5 1.2 

NOE2 13 142.6 1.2 

NOE2 14 142.6 1.2 

NOE2 15 142.7 1.2 

NOE2 16 142.7 1.2 

NOE2 17 142.7 1.2 

NOE2 18 142.7 1.2 

NOE2 19 142.7 1.2 

NOE2 20 142.7 1.2 

NOE2 21 142.7 1.2 

NOE2 22 142.7 1.2 

NOE2 23 142.7 1.2 

NOE2 24 142.7 1.2 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-40 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

NOE3 -6 0.0 0.0 

NOE3 -5 0.3 0.0 

NOE3 -4 8.0 0.0 

NOE3 -3 39.6 0.1 

NOE3 -2 189.0 1.6 

NOE3 -1 365.8 3.6 

NOE3 0 399.8 4.5 

NOE3 1 416.2 5.6 

NOE3 2 510.1 22.3 

NOE3 3 528.7 25.0 

NOE3 4 563.7 31.5 

NOE3 5 577.3 33.3 

NOE3 6 594.1 33.8 

NOE3 7 618.7 34.4 

NOE3 8 629.6 35.2 

NOE3 9 636.6 35.9 

NOE3 10 649.1 37.6 

NOE3 11 666.7 40.3 

NOE3 12 668.4 40.4 

NOE3 13 671.0 40.9 

NOE3 14 677.1 42.2 

NOE3 15 678.8 42.5 

NOE3 16 680.2 42.8 

NOE3 17 680.3 42.8 

NOE3 18 680.3 42.8 

NOE3 19 680.3 42.8 

NOE3 20 680.3 42.8 

NOE3 21 680.3 42.8 

NOE3 22 680.3 42.8 

NOE3 23 680.3 42.8 

NOE3 24 680.3 42.8 

NOE3 25 680.3 42.8 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-41 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

NOE4 -1 0.0 0.0 

NOE4 0 0.1 0.1 

NOE4 1 24.9 12.7 

NOE4 2 32.6 19.4 

NOE4 3 34.6 20.7 

NOE4 4 49.1 33.9 

NOE4 5 54.3 38.7 

NOE4 6 55.7 40.0 

NOE4 7 56.6 40.7 

NOE4 8 57.4 41.2 

NOE4 9 58.0 41.7 

NOE4 10 59.5 42.9 

NOE4 11 60.1 43.3 

NOE4 12 60.5 43.7 

NOE4 13 61.9 44.4 

NOE4 14 62.9 45.1 

NOE4 15 63.6 45.6 

NOE4 16 65.4 47.4 

NOE4 17 65.8 47.7 

NOE4 18 65.9 47.9 

NOE4 19 66.0 47.9 

NOE4 20 66.0 47.9 

NOE4 21 66.0 47.9 

NOE4 22 66.0 47.9 

NOE4 23 66.0 47.9 

NOE4 24 66.0 47.9 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-42 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

NOE5 -12 0.0 0.0 

NOE5 -11 0.1 0.0 

NOE5 -10 1.2 0.0 

NOE5 -9 5.2 0.0 

NOE5 -8 42.7 0.0 

NOE5 -7 241.5 1.8 

NOE5 -6 962.3 9.9 

NOE5 -5 2,316.5 25.4 

NOE5 -4 3,083.4 42.5 

NOE5 -3 3,484.4 63.4 

NOE5 -2 4,121.0 146.5 

NOE5 -1 4,560.7 225.3 

NOE5 0 4,940.0 325.0 

NOE5 1 5,066.5 359.3 

NOE5 2 5,183.0 378.0 

NOE5 3 5,430.1 467.7 

NOE5 4 5,594.9 508.4 

NOE5 5 5,674.0 517.9 

NOE5 6 5,778.1 529.9 

NOE5 7 5,878.2 540.8 

NOE5 8 5,945.3 553.9 

NOE5 9 5,974.5 563.8 

NOE5 10 5,993.7 569.3 

NOE5 11 6,007.4 575.7 

NOE5 12 6,021.6 584.6 

NOE5 13 6,025.7 587.1 

NOE5 14 6,028.0 589.0 

NOE5 15 6,034.6 595.2 

NOE5 16 6,036.7 597.1 

NOE5 17 6,037.3 597.7 

NOE5 18 6,037.5 597.9 

NOE5 19 6,037.5 597.9 

NOE5 20 6,037.5 597.9 

NOE5 21 6,037.5 597.9 

NOE5 22 6,037.5 597.9 

NOE5 23 6,037.5 597.9 

NOE5 24 6,037.5 597.9 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-43 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

OM1 -12 0.0 0.0 

OM1 -11 0.0 0.0 

OM1 -10 0.0 0.0 

OM1 -9 0.0 0.0 

OM1 -8 0.0 0.0 

OM1 -7 8.4 0.0 

OM1 -6 102.8 0.0 

OM1 -5 401.7 0.0 

OM1 -4 782.6 0.0 

OM1 -3 1,008.1 0.7 

OM1 -2 1,269.7 3.3 

OM1 -1 1,531.3 9.3 

OM1 0 1,713.2 16.7 

OM1 1 1,864.2 29.1 

OM1 2 2,012.0 51.5 

OM1 3 2,208.9 89.3 

OM1 4 2,339.4 131.2 

OM1 5 2,458.6 188.1 

OM1 6 2,526.4 221.6 

OM1 7 2,574.8 234.9 

OM1 8 2,603.0 238.7 

OM1 9 2,625.3 242.1 

OM1 10 2,650.3 245.8 

OM1 11 2,675.0 251.1 

OM1 12 2,689.4 255.4 

OM1 13 2,702.4 261.0 

OM1 14 2,717.7 268.7 

OM1 15 2,725.8 272.8 

OM1 16 2,731.7 276.2 

OM1 17 2,737.6 279.5 

OM1 18 2,738.9 280.3 

OM1 19 2,739.7 280.9 

OM1 20 2,739.9 281.1 

OM1 21 2,739.9 281.1 

OM1 22 2,739.9 281.1 

OM1 23 2,739.9 281.1 

OM1 24 2,739.9 281.1 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-44 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

OM2 -11 0.0 0.0 

OM2 -10 0.0 0.0 

OM2 -9 0.0 0.0 

OM2 -8 0.0 0.0 

OM2 -7 2.0 0.0 

OM2 -6 69.9 0.0 

OM2 -5 347.6 0.0 

OM2 -4 677.1 0.0 

OM2 -3 884.1 0.2 

OM2 -2 1,040.1 1.8 

OM2 -1 1,196.3 4.1 

OM2 0 1,260.9 6.9 

OM2 1 1,376.7 26.7 

OM2 2 1,482.4 44.3 

OM2 3 1,573.3 64.0 

OM2 4 1,685.3 103.2 

OM2 5 1,763.0 142.9 

OM2 6 1,835.0 188.5 

OM2 7 1,858.6 191.5 

OM2 8 1,877.2 194.5 

OM2 9 1,893.8 197.9 

OM2 10 1,909.8 201.1 

OM2 11 1,922.1 203.6 

OM2 12 1,930.7 207.6 

OM2 13 1,938.9 212.8 

OM2 14 1,945.5 218.0 

OM2 15 1,948.6 220.5 

OM2 16 1,949.5 221.2 

OM2 17 1,950.0 221.5 

OM2 18 1,950.9 222.4 

OM2 19 1,951.1 222.5 

OM2 20 1,951.1 222.6 

OM2 21 1,951.1 222.6 

OM2 22 1,951.1 222.6 

OM2 23 1,951.1 222.6 

OM2 24 1,951.1 222.6 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-45 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

OM3 -10 0.0 0.0 

OM3 -9 0.0 0.0 

OM3 -8 0.0 0.0 

OM3 -7 0.0 0.0 

OM3 -6 0.8 0.0 

OM3 -5 8.7 0.0 

OM3 -4 29.8 0.0 

OM3 -3 65.9 0.1 

OM3 -2 154.4 1.1 

OM3 -1 376.1 4.7 

OM3 0 661.8 16.0 

OM3 1 1,030.8 40.5 

OM3 2 1,364.6 87.0 

OM3 3 1,708.7 166.7 

OM3 4 2,030.1 287.9 

OM3 5 2,281.0 456.3 

OM3 6 2,436.6 582.1 

OM3 7 2,581.6 691.5 

OM3 8 2,696.6 774.4 

OM3 9 2,793.1 832.9 

OM3 10 2,859.7 866.2 

OM3 11 2,910.0 887.7 

OM3 12 2,952.5 903.5 

OM3 13 2,984.3 919.7 

OM3 14 3,010.0 936.8 

OM3 15 3,029.5 950.5 

OM3 16 3,045.5 963.0 

OM3 17 3,060.7 976.3 

OM3 18 3,069.1 984.1 

OM3 19 3,073.2 987.9 

OM3 20 3,075.8 990.3 

OM3 21 3,077.0 991.5 

OM3 22 3,077.3 991.8 

OM3 23 3,077.7 992.1 

OM3 24 3,077.7 992.2 

OM3 25 3,077.7 992.2 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-46 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

OM4 -6 0.0 0.0 

OM4 -5 2.9 0.0 

OM4 -4 5.8 0.0 

OM4 -3 8.8 0.1 

OM4 -2 27.7 0.7 

OM4 -1 68.5 1.8 

OM4 0 142.1 7.2 

OM4 1 219.1 13.7 

OM4 2 308.1 28.6 

OM4 3 420.5 59.1 

OM4 4 618.1 158.1 

OM4 5 790.4 288.4 

OM4 6 880.0 366.3 

OM4 7 954.5 430.0 

OM4 8 982.0 449.4 

OM4 9 1,013.4 470.1 

OM4 10 1,034.5 480.0 

OM4 11 1,055.3 490.9 

OM4 12 1,075.7 500.7 

OM4 13 1,096.4 514.2 

OM4 14 1,110.8 525.7 

OM4 15 1,119.4 531.8 

OM4 16 1,131.1 539.9 

OM4 17 1,139.5 547.2 

OM4 18 1,145.6 552.7 

OM4 19 1,151.2 557.9 

OM4 20 1,153.5 560.1 

OM4 21 1,155.0 561.6 

OM4 22 1,155.3 561.9 

OM4 23 1,155.6 562.1 

OM4 24 1,155.6 562.2 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-47 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

OM5 -9 0.0 0.0 

OM5 -8 0.0 0.0 

OM5 -7 0.0 0.0 

OM5 -6 0.7 0.0 

OM5 -5 0.8 0.0 

OM5 -4 6.8 0.1 

OM5 -3 37.9 0.4 

OM5 -2 200.5 2.3 

OM5 -1 785.0 10.6 

OM5 0 1,483.2 34.8 

OM5 1 2,167.0 79.3 

OM5 2 2,859.8 177.1 

OM5 3 3,721.4 358.5 

OM5 4 4,837.3 827.4 

OM5 5 5,522.0 1,296.7 

OM5 6 6,034.1 1,701.3 

OM5 7 6,834.7 2,387.0 

OM5 8 7,538.2 2,966.8 

OM5 9 8,112.3 3,415.2 

OM5 10 8,574.3 3,758.4 

OM5 11 8,920.2 4,012.9 

OM5 12 9,305.7 4,303.0 

OM5 13 9,511.7 4,449.5 

OM5 14 9,679.7 4,572.3 

OM5 15 9,846.8 4,675.5 

OM5 16 10,031.8 4,800.2 

OM5 17 10,144.1 4,897.0 

OM5 18 10,260.3 5,003.6 

OM5 19 10,353.9 5,092.2 

OM5 20 10,401.1 5,137.0 

OM5 21 10,436.4 5,170.6 

OM5 22 10,457.9 5,191.1 

OM5 23 10,488.4 5,220.1 

OM5 24 10,512.5 5,242.9 

OM5 25 10,517.8 5,247.9 

OM5 26 10,520.0 5,250.0 

OM5 27 10,520.5 5,250.5 

OM5 28 10,520.6 5,250.6 

OM5 29 10,520.6 5,250.6 

OM5 30 10,520.6 5,250.6 

OM5 31 10,520.6 5,250.6 

OM5 32 10,520.6 5,250.6 

SB1 -5 0.0 0.0 

SB1 -4 0.3 0.0 

SB1 -3 8.7 0.0 

SB1 -2 53.7 0.0 

SB1 -1 196.4 0.0 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-48 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

SB1 0 476.7 0.3 

SB1 1 876.7 4.7 

SB1 2 1,262.7 8.3 

SB1 3 1,496.9 11.5 

SB1 4 1,722.8 17.3 

SB1 5 1,983.3 32.6 

SB1 6 2,159.9 53.1 

SB1 7 2,306.0 78.2 

SB1 8 2,403.4 98.9 

SB1 9 2,480.3 116.1 

SB1 10 2,542.2 124.8 

SB1 11 2,582.0 129.2 

SB1 12 2,617.0 133.5 

SB1 13 2,650.5 138.2 

SB1 14 2,676.6 141.4 

SB1 15 2,694.2 144.2 

SB1 16 2,707.6 147.1 

SB1 17 2,717.6 149.3 

SB1 18 2,725.5 150.9 

SB1 19 2,729.7 152.2 

SB1 20 2,731.4 152.8 

SB1 21 2,732.3 153.3 

SB1 22 2,732.8 153.7 

SB1 23 2,733.3 154.2 

SB1 24 2,733.3 154.2 

SB1 25 2,733.4 154.3 

SB1 26 2,733.5 154.4 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-49 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (continued) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

SB3 -5 0.0 0.0 

SB3 -4 0.0 0.0 

SB3 -3 0.1 0.0 

SB3 -2 0.3 0.0 

SB3 -1 3.7 0.0 

SB3 0 70.6 0.0 

SB3 1 217.6 0.3 

SB3 2 633.5 3.7 

SB3 3 1,149.7 10.4 

SB3 4 1,524.6 23.5 

SB3 5 1,723.6 33.4 

SB3 6 1,836.2 44.2 

SB3 7 1,938.8 63.5 

SB3 8 2,077.3 117.4 

SB3 9 2,190.4 185.2 

SB3 10 2,298.0 239.2 

SB3 11 2,393.5 270.9 

SB3 12 2,455.8 286.1 

SB3 13 2,490.8 294.3 

SB3 14 2,509.6 297.9 

SB3 15 2,524.3 301.4 

SB3 16 2,540.0 306.9 

SB3 17 2,554.3 314.6 

SB3 18 2,563.1 320.5 

SB3 19 2,568.0 323.9 

SB3 20 2,573.5 327.3 

SB3 21 2,576.7 329.7 

SB3 22 2,578.6 331.2 

SB3 23 2,580.0 332.6 

SB3 24 2,580.4 333.0 

SB3 25 2,580.6 333.1 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-50 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 12-5: Stage-Damage for All Sub-basins (Concluded) 

Subbasin Name Water Elevation Pre-Katrina Stage-Damage Post-Katrina Stage-Damage June 2006 

SB4 1 0.0 0.0 

SB4 2 0.4 0.0 

SB4 3 7.3 0.1 

SB4 4 32.7 1.3 

SB4 5 89.7 6.4 

SB4 6 190.2 21.7 

SB4 7 295.3 48.0 

SB4 8 387.5 81.1 

SB4 9 438.6 100.6 

SB4 10 465.4 106.5 

SB4 11 491.2 112.5 

SB4 12 497.1 113.9 

SB4 13 503.4 114.8 

SB4 14 515.7 116.6 

SB4 15 527.7 119.1 

SB4 16 537.8 122.5 

SB4 17 544.7 124.8 

SB4 18 549.4 126.0 

SB4 19 552.6 126.7 

SB4 20 553.7 127.2 

SB4 21 554.3 127.5 

SB4 22 554.5 127.5 

SB4 23 554.5 127.6 

SB4 24 554.5 127.6 

SB4 25 554.5 127.6 

SB4 26 554.5 127.6 

SB4 27 554.5 127.6 

SB4 28 554.5 127.6 

SB4 29 554.5 127.6 

SB4 30 554.5 127.6 

SB4 31 554.5 127.6 

SB4 32 554.5 127.6 

SB4 33 554.5 127.6 

SB4 34 554.5 127.6 

SB4 35 554.5 127.6 

SB4 36 554.5 127.6 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-12-51 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

References 

Kevin McCarthy, D.J. Peterson, Narayan Sastry, and Michael Pollard. (2006) “The Repopulation of New 

Orleans after Hurricane Katrina.” RAND Gulf States Policy Institute. RAND URL: 

http://www.rand.org/

Aboelata, M. and D.S. Bowles. 2005. LIFESim: A Model for Estimating Dam Failure Life Loss. Draft 

report to the Institute for Water Resources, US Army Corps of Engineers by the Institute for Dam 

Safety Risk Management, Utah State University, Logan, Utah. 

Aboelata, M., D.S. Bowles, and D.M. McClelland. 2003. ‘Life-loss Estimation for Floods including Dam 

Failure.’ GIS Model for Estimating Dam Failure Life Loss. In Y.Y. Haimes and D.A. Moser, (eds.), 

American Society of Civil Engineers. 

Aboelata, M., D.S. Bowles, and D.M. McClelland. 2004a. ‘A Model for Estimating Dam Failure Life 

Loss.’ ANCOLD Bulletin 127:43-62. August. 

Aboelata, M., D.S. Bowles, and A. Chen. 2004b. ‘Transportation model for evacuation in estimating dam 

failure life loss.’ Proceedings of the Australian National Committee on Large Dams Conference, 

Melbourne, Victoria, Australia. 

Abt Associates, Inc. 2006. Estimating Loss of Life from Hurricane-Related Flooding in the Greater New 

Orleans Area. Loss-of-Life Modeling Report. Prepared for U.S. Army Corps of Engineers, Institute 

for Water Resources, Alexandria, VA. 

McClelland, D.M. 2000. Personal Communication of Updated Probability Distributions for Fatality Rates 

for Each Flood Zone. November 30. 

McClelland, D. M. 2002. “A team approach to improving dam-safety emergency action plans, flood 

maps, and emergency response plans using life-loss risk assessment.” EAP 2002: International 

Workshop for Emergency Preparedness at Dams, sponsored by FERC and ASDSO. 

McClelland, D.M. and Bowles, D.S. 2002. ‘Estimating Life Loss for Dam Safety Risk Assessment - a 

Review and New Approach.’ Institute for Water Resources, U.S. Army Corps of Engineers, 

Alexandria, VA. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-1 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 13 

Risk Analysis Results 

Introduction 

The results of the risk analysis of the New Orleans Hurricane Protection System (HPS) are 

presented in this appendix. The purpose of the analysis was to evaluate the risk of inundation and 

the corresponding risks to life and property posed by the HPS prior to the arrival of hurricane 

Katrina and as it existed on 1 June 2007. These points in time studied are referred to as the “PreKatrina (or Pre-K)” and “Current” conditions, respectively. The spreadsheet computer program 

FoRTE (presented in Appendix 15) developed by the risk team was used as the basis for 

establishing elevation-exceedence curves for these conditions for three states of pumping system 

effectiveness. Analyses were run for no pumping, pumping at 50% of capacity and at 100% of 

capacity. The computer runs included pumping by modifying the sub-basin rainfall by the 

amount of pumping volume expected for each storm. Pumping volumes were estimated 

deterministically and subtracted from the rainfall volume calculated for each storm. 

Elevations used to produce inundation mapping were selected using several factors to modify 

the computer generated elevation-exceedence curves. The computer runs did not include 

overtopping water volumes due to wave run-up so this item was added to the runs using 

deterministic models. The risk program conducted interflow analyses between sub-basins at the 

basin level by individual storms but did not do interflow analyses for the basins when the storms 

were aggregated. Therefore, the program results at the 50-, 100-, and 500-year exceedence rates 

were examined and balanced by looking at the water volumes produced at each exceedence level 

using the stage-storage relationships for the sub-basins. If the interflow elevations between 

sub-basins were exceeded, water volumes were redistributed and new water surface elevations 

determined. Note that the exceedence rates are conditional on the storm set provided to the risk 

team with frequencies as shown in Appendix 8, which do not consider tropical storms and lower 

intensity, more frequent hurricanes. 

Measuring System Performance 

The primary measure used to evaluate the risks associated with the New Orleans HPS is 

water elevations within the system based on the amount of inflow water volume from 

overtopping, breaching, rainfall and closure structures left open. These elevations were used to 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-2 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

determine the rate of overtopping of the reaches, construct inundation maps, and to determine 

life and economic risks.

Overtopping Rates 

One measure of the effectiveness of the HPS is how often the system can be expected to be 

overtopped by the storms predicted to strike the New Orleans area. This measure, while not risk 

based, provided decision makers with additional information about system performance. The 

rates of overtopping were determined for the Pre-K and Current HPS for the total 152 storm set 

provided by the IPET/LaCPR storm team. Note that while the entire 152 storm set was used to 

determine overtopping rates, only the 77 storms with frequencies provided by the storm team 

were used in the actual risk analysis. The number of times that each reach within the basins was 

overtopped was calculated by counting the storm events whose peak surge exceeded the reach 

elevation. The overtopping rate was then determined by dividing the total number of times that a 

reach was overtopped by the number of storms (152). Basin and HPS overtopping rates were 

also determined and are presented in Table 13-1. The rates are also graphically shown in 

Figures 13-10 and 13-11 for the Pre-K and Current HPS, respectively, with color coding to rank 

areas by the overtopping rate. 

As shown in Table 13-1, the number of overtopping events in the Current HPS is reduced 

somewhat from those of Pre-K levels in most basins and at many reaches. This is due to 

restoring damaged levees to the Pre-Katrina authorized elevation, which in many cases is higher 

than the actual Pre-K elevation because of subsidence and datum changes. Areas that show high 

rates of overtopping, greater than 25% of the storms, in the Pre-Katrina HPS are in New Orleans 

East along the IHNC and MRGO, all of the southern portions of Jefferson West Bank and most 

of Plaquemines. The overall rates of overtopping, however, are very similar in both HPS. 

Table 13-1 Overtopping Rates for Reaches and Sub-basins 

Pre-K Current 

Station 

No.

Reach 

Name

Subbasin No of 

OT 

% of storms 

OT Color

No of 

OT 

% of storms 

OT Color Color Legend

1 NOE 1 NOE5 5.0 0.033 6.0 0.039 0=Dark Green

2 NOE 2 NOE5 5.0 0.033 6.0 0.039 0-.1 = 

Lt Green 

3 NOE 3 NOE5 5.0 0.033 6.0 0.039 .1-.2 = Lt Blue

4 NOE 4 NOE5 5.0 0.033 6.0 0.039 .2-.3 = Yellow 

5 NOE 5 NOE5 5.0 0.033 6.0 0.039 .3-.4 = 

Orange 

6 NOE 6 NOE5 1.0 0.007 2.0 0.013 .4-.5 = Red 

7 NOE 7 NOE5 1.0 0.007 2.0 0.013 

8 NOE 8 NOE5 2.0 0.013 3.0 0.020 

9 NOE 9 NOE1 0.0 0.000 0.0 0.000 

10 NOE 10 NOE1 2.0 0.013 3.0 0.020 

11 NOE 11 NOE1 3.0 0.020 4.0 0.026 

12 NOE 12 NOE1 3.0 0.020 4.0 0.026 

13 NOE 13 NOE1 3.0 0.020 4.0 0.026 

14 NOE 14 NOE1 13.0 0.086 14.0 0.092 

15 NOE 15 NOE1 18.0 0.118 10.0 0.066 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-3 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-1 Overtopping Rates for Reaches and Sub-basins 

Pre-K Current 

Station 

No.

Reach 

Name

Subbasin No of 

OT 

% of storms 

OT Color

No of 

OT 

% of storms 

OT Color Color Legend

16 NOE 16 NOE1 27.0 0.178 14.0 0.092 

17 NOE 17 NOE2 34.0 0.224 18.0 0.118 

18 NOE 18 NOE2 27.0 0.178 27.0 0.178 

19 NOE 19 NOE3 28.0 0.184 28.0 0.184 

20 NOE 20 NOE3 39.0 0.257 38.0 0.250 

21 NOE 21 NOE3 39.0 0.257 38.0 0.250 

22 NOE 22 NOE3 34.0 0.224 33.0 0.217 

23 NOE 23 NOE4 56.0 0.368 55.0 0.362 

24 NOE 24 NOE4 56.0 0.368 55.0 0.362 

25 NOE 25 NOE4 54.0 0.355 53.0 0.349 

26 NOE 26 NOE4 52.0 0.342 51.0 0.336 

27 NOE 27 NOE4 51.0 0.336 50.0 0.329 

28 NOE 28 NOE4 69.0 0.454 54.0 0.355 

29 NOE 29 NOE5 3.0 0.020 4.0 0.026 

Total NOE 640.0 0.145 594.0 0.135 

30 JE1 JE3 12.0 0.079 13.0 0.086 

31 JE2 JE3 6.0 0.039 7.0 0.046 

32 JE3 JE3 1.0 0.007 2.0 0.013 

33 JE4 JE3 0.0 0.000 0.0 0.000 

34 JE5 JE2 0.0 0.000 0.0 0.000 

35 JE6 JE2 0.0 0.000 0.0 0.000 

37 JE8 JE1 0.0 0.000 0.0 0.000 

38 JE9 JE3 38.0 0.250 38.0 0.250 

Total JE 57.0 0.013 60.0 0.014 

39 SC1 SC1 32.0 0.211 31.0 0.204 

40 SC2 SC1 44.0 0.289 17.0 0.112 

41 SC3 SC1 37.0 0.243 36.0 0.237 

42 SC4 SC2 0.0 0.000 0.0 0.000 

43 SC5 SC2 0.0 0.000 0.0 0.000 

44 SC6 SC2 0.0 0.000 0.0 0.000 

Total SC 113.0 0.124 84.0 0.092 

45 OM1 OM2 0.0 0.000 0.0 0.000 

46 OM2 OM2 0.0 0.000 0.0 0.000 

47 OM3 OM2 0.0 0.000 0.0 0.000 

48 OM4 OM2 0.0 0.000 0.0 0.000 

49 OM5 OM2 0.0 0.000 0.0 0.000 

50 OM6 OM1 0.0 0.000 0.0 0.000 

51 OM7 OM1 0.0 0.000 0.0 0.000 

52 OM8 OM1 0.0 0.000 0.0 0.000 

53 OM9 OM1 0.0 0.000 0.0 0.000 

54 OM10 OM2 3.0 0.020 4.0 0.026 

55 OM11 OM2 0.0 0.000 0.0 0.000 

56 OM12 OM1 0.0 0.000 0.0 0.000 

57 OM13 OM1 0.0 0.000 0.0 0.000 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-4 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-1 Overtopping Rates for Reaches and Sub-basins 

Pre-K Current 

Station 

No.

Reach 

Name

Subbasin No of 

OT 

% of storms 

OT Color

No of 

OT 

% of storms 

OT Color Color Legend

58 OM14 OM1 0.0 0.000 0.0 0.000 

59 OM15 OM1 0.0 0.000 0.0 0.000 

60 OM16 OM1 0.0 0.000 0.0 0.000 

61 OM17 OM1 2.0 0.013 3.0 0.020 

62 OM18 OM3 33.0 0.217 32.0 0.211 

63 OM19 OM3 49.0 0.322 20.0 0.132 

64 OM20 OM3 51.0 0.336 50.0 0.329 

65 OM21 OM3 0.0 0.000 0.0 0.000 

66 OM22 OM3 0.0 0.000 0.0 0.000 

67 OM23 OM3 0.0 0.000 0.0 0.000 

68 OM24 OM5 0.0 0.000 0.0 0.000 

69 OM25 OM5 0.0 0.000 0.0 0.000 

70 OM26 OM5 0.0 0.000 0.0 0.000 

71 OM27 OM4 0.0 0.000 0.0 0.000 

36 JE7 OM4 0.0 0.000 0.0 0.000 

 Total OM 138.0 0.032 109.0 0.026 

72 SB1 SB1 50.0 0.329 20.0 0.132 

73 SB2 SB2 12.0 0.079 13.0 0.086 

74 SB3 SB2 20.0 0.132 21.0 0.138 

75 SB4 SB5 2.0 0.013 0.0 0.000 

76 SB5 SB4 0.0 0.000 0.0 0.000 

77 SB6 SB4 0.0 0.000 0.0 0.000 

78 SB7 SB3 0.0 0.000 0.0 0.000 

79 SB8 SB1 0.0 0.000 0.0 0.000 

80 SB9 SB1 0.0 0.000 0.0 0.000 

Total SB 84.0 0.061 54.0 0.039 

81 PL1 PL1 140.0 0.921 133.0 0.875 

82 PL2 PL1 48.0 0.316 48.0 0.316 

83 PL3 PL1 0.0 0.000 0.0 0.000 

84 PL4 PL2 39.0 0.257 39.0 0.257 

85 PL5 PL2 2.0 0.013 3.0 0.020 

86 PL6 PL3 136.0 0.895 130.0 0.855 

87 PL7 PL3 0.0 0.000 0.0 0.000 

88 PL8 PL7 31.0 0.204 30.0 0.197 

89 PL9 PL7 3.0 0.020 4.0 0.026 

90 PL10 PL7 20.0 0.132 21.0 0.138 

91 PL11 PL8 21.0 0.138 22.0 0.145 

92 PL12 PL8 1.0 0.007 2.0 0.013 

94 PL13 PL8 44.0 0.289 45.0 0.296 

95 PL14 PL8 2.0 0.013 3.0 0.020 

96 PL15 PL8 0.0 0.000 0.0 0.000 

96 PL16 PL9 13.0 0.086 14.0 0.092 

97 PL17 PL10 22.0 0.145 23.0 0.151 

98 PL18 PL10 26.0 0.171 27.0 0.178 

99 PL19 PL9 44.0 0.289 45.0 0.296 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-5 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-1 Overtopping Rates for Reaches and Sub-basins 

Pre-K Current 

Station 

No.

Reach 

Name

Subbasin No of 

OT 

% of storms 

OT Color

No of 

OT 

% of storms 

OT Color Color Legend

100 PL20 PL4 0.0 0.000 0.0 0.000 

101 PL21 PL5 0.0 0.000 0.0 0.000 

102 PL22 PL4 0.0 0.000 0.0 0.000 

103 PL23 PL5 0.0 0.000 0.0 0.000 

104 PL24 PL1 3.0 0.020 4.0 0.026 

105 PL25 PL6 1.0 0.007 2.0 0.013 

106 PL26 PL1 0.0 0.000 0.0 0.000 

107 PL27 PL6 0.0 0.000 0.0 0.000 

Total PL 596.0 0.145 595.0 0.145 

108 CW1 JW1 133.0 0.875 128.0 0.842 

109 CW2 JW2 122.0 0.803 118.0 0.776 

110 CW3 JW2 127.0 0.836 115.0 0.757 

111 CW4 JW2 127.0 0.836 122.0 0.803 

112 CW5 JW2 136.0 0.895 130.0 0.855 

113 CW6 JW2 115.0 0.757 111.0 0.730 

114 CW7 JW3 26.0 0.171 26.0 0.171 

115 CW8 JW1 0.0 0.000 0.0 0.000 

116 WH1 JW3 19.0 0.125 19.0 0.125 

117 WH2 JW3 97.0 0.638 68.0 0.447 

118 WH3 JW3 48.0 0.316 47.0 0.309 

119 WH4 JW3 1.0 0.007 2.0 0.013 

120 WH5 JW3 4.0 0.026 2.0 0.013 

121 WH6 JW3 26.0 0.171 26.0 0.171 

122 WH7 JW3 96.0 0.632 71.0 0.467 

123 WH8 JW3 48.0 0.316 6.0 0.039 

124 WH9 JW3 0.0 0.000 0.0 0.000 

125 WH10 JW3 0.0 0.000 0.0 0.000 

127 HA2 JW4 0.0 0.000 0.0 0.000 

128 HA3 JW4 122.0 0.803 47.0 0.309 

129 HA4 JW4 60.0 0.395 59.0 0.388 

135 HA10 JW4 0.0 0.000 0.0 0.000 

Total JW 1307.0 0.391 1097.0 0.328 

126 HA1 OW2 0.0 0.000 0.0 0.000 

130 HA5 OW2 17.0 0.112 18.0 0.118 

131 HA6 OW1 0.0 0.000 0.0 0.000 

132 HA7 OW1 3.0 0.020 4.0 0.026 

133 HA8 OW1 145.0 0.954 138.0 0.908 

134 HA9 OW2 94.0 0.618 91.0 0.599 

Total OW 259.0 0.284 251.0 0.275 

Total All Basins 3194.00 0.157 2844.00 0.140 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-6 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Expected Water Surface Elevations 

The results of the risk analysis are presented in terms of the water surface elevations 

expected in each sub-basin for the “no pumping,” “50% pumping,” and “100% pumping” 

scenarios in Tables 13-2 to 13-4 and the three exceedence (50, 100 and 500-Year) rates used to 

construct the inundation maps. Inundation maps are provided in the figures at the end of this 

Appendix.

Results without pumping – Table 13-2 shows the water elevations within each sub-basin 

for the Pre-K and Current HPS if the pumping system is considered to not be operating. This 

pumping scenario most closely approximates the actual performance of the system during 

hurricanes with intensities and frequencies similar to Katrina. The 50-year water volumes and 

resulting elevations (Figures 13-2 and 13-3) are primarily hurricane rainfall in the basins, while 

the 100-year elevations (Figures 13-4, 13-5 and 13-6) show the impact of the large amount of 

overtopping, levee or wall breaching and open closure structures. 

Reductions in water surface elevations expected for the Current HPS were noted in the OM 

and SB basins, but were largely unchanged in all other basins. SB showed the most reduction in 

water elevation due to the improvements made in the Current HPS with reductions of 2 to 2.5 

feet. All basins fill during 500-year events in both the Pre-Katrina and Current HPS 

(Figures 13-7, 13-8 and 13-9). 

Results with 50% pumping – Table 13-3 shows the water elevations within each sub-basin 

for the Pre-K and Current HPS if the pumping system is considered to be operating at 50% 

capacity. The 50-year water inflow volumes and resulting elevations are primarily hurricane 

rainfall in the basins. This pumping scenario most closely approximates the possible 

performance of the system during a 50-year event for both Pre-K and Current HPS when the 

pumps would only have to evacuate hurricane rains, which would be expected to be less than the 

design rainfall for the pumps (about a 10-year tropical storm). Therefore, the differences in water 

elevations between the no pumping and 50% pumping scenarios, for both the Pre-K and Current 

HPS, are small for the 50-year event. 

Results with 100% pumping – Table 13-4 shows the water elevations within each sub-basin 

for the Pre-K and Current HPS if the pumping system is considered to be operating at 100% of 

capacity. This scenario is presented for comparison purposes with the other scenarios only since 

the New Orleans pumping system, or any other pumping system, cannot realistically be expected 

to perform at this level. The significant impact is the reduction of interior flooding. Water 

elevations could be reduced by as much as 4 feet in show areas for the 50 year event, while the 

100-Year show the impact of the large amount of overtopping, levee or wall breaching and open 

closure structures. During hurricanes with intensities and frequencies similar to Katrina, the PreK pumping system would be expected to perform as shown, as it did during Katrina, and not be 

able to handle flooding from overtopping or breaching inflows. Reductions in water surface 

elevations expected for the Current HPS were noted in the OM and SB basins, but were largely 

unchanged in all other basins. All basins fill during 500-year events in both the Pre-Katrina and 

Current HPS. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-7 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-2 Risk Analysis Results Without Pumping 

Elevations NAVD88 2004.65 

50-year elevations 100-year elevations 500-year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 -1 -1 1 1 6 6 

OW2 -3 -3 4 4 8 8 

NOE1 0 0 2 1 12 13 

NOE2 -4 -4 2 2 12 13 

NOE3 -4 -4 2 2 12 13 

NOE4 -1 -1 4 4 12 13 

NOE5 -8 -8 -1 -1 12 13 

OM1 -5 -5 3 3 14 14 

OM2 -5 -5 3 -2 14 14 

OM3 -1 -1 3 3 14 14 

OM4 -1 -1 3 -2 14 14 

OM5 -1 -1 3 2 14 14 

SB1 -1 -1 12 10 14 14 

SB2 1 1 12 10 14 14 

SB3 0 0 12 10 14 14 

SB4 2 2 12 10 14 14 

SB5 3 3 12 10 14 14 

JE1 3 3 4 4 14 14 

JE2 -4 -4 -3 -3 14 14 

JE3 -5 -5 -3 -3 1 14 

JW1 0 0 4 4 8 8 

JW2 -4 -4 4 4 8 8 

JW3 -2 -2 4 4 8 8 

JW4 -5 -5 4 4 8 8 

PL11 -2 -2 0 -1 6 9 

SC1 2 2 4 5 10 10 

SC2 4 4 5 5 10 10 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 13-1 – 50-Year Inundation Elevations for Pre-K and Current HPS Figure 13-2 – 50-Year Inundation Elevations for Pre-K HPS with Pumping Scenarios 

50 Year Elevations

-10

-8-6-4-2

0246

OW1OW2NOE1NOE2NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1

Subbasin

Elevation (ft) 

Pre- Katrina 50 Year Elevations

-14-12-10

-8-6-4-2

0246

OW1OW2 NOE1 NOE2 NOE3 NOE4 NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1SC2

Subbasin

Elevation (ft) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 13-3 – 50-Year Inundation Elevations for Current HPS with Pumping Scenarios Figure 13-4 – 100-Year Inundation Elevations for Pre-K and Current HPS 

Current 50 Year Elevations

-14-12-10

-8-6-4-2

0246

OW1OW2 NOE1 NOE2NOE3 NOE4 NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1SC2

Subbasin

Elevation (ft) 

100 Year Elevations

-4-2

02468

101214

OW1OW2NOE1NOE2NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1S

Subbasin

Elevation (ft) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 13-5 – 100-Year Inundation Elevations for Pre-K HPS with Pumping Scenarios Figure 13-6 – 100-Year Inundation Elevations for Current HPS with Pumping Scenarios 

Pre- Katrina 100 Year Elevations

-15-10

-5

05

1015

OW1OW2NOE1NOE2 NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1SC2

Subbasin

Elevation (ft) 

 Current 100 Year Elevations

-15-10

-5

05

1015

OW1OW2NOE1 NOE2NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1SC2

Subbasin

Elevation (ft) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 13-7 – 500-Year Inundation Elevations for Pre-K and Current HPS Figure 13-8 – 500-Year Inundation Elevations for Pre-K with Pumping Scenarios 

500 Year Elevations

02468

10121416

OW1OW2NOE1NOE2NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1S

Subbasin

Elevation (ft) 

 Pre-Katrina 500 Year Elevations

02468

10121416

OW1OW2NOE1NOE2NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1SC2

Subbasin

Elevation (ft) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 13-9 – 500-Year Inundation Elevations for Current HPS with Pumping Scenarios 

Current 500 Year Elevations

02468

10121416

OW1OW2NOE1NOE2NOE3NOE4NOE5OM1OM2OM3OM4OM5SB1SB2SB3SB4SB5JE1JE2JE3JW1JW2JW3JW4PL11SC1SC2

Subbasin

Elevation (ft) 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-13 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-3 Risk Analysis Results With 50% Pumping 

Elevations NAVD88 2004.65 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 -1 -1 0 0 6 6 

OW2 -3 -3 3 3 8 8 

NOE1 0 0 2 1 12 13 

NOE2 -5 -5 1 1 11 12 

NOE3 -5 -5 1 1 11 12 

NOE4 -2 -2 2 2 11 12 

NOE5 -9 -9 -2 -2 11 12 

OM1 -7 -7 2 1 12 12 

OM2 -12 -12 1 -7 11 11 

OM3 -6 -6 2 1 12 12 

OM4 -1 -1 3 -2 13 13 

OM5 -4 -4 2 0 12 12 

SB1 -5 -5 11 8 12 12 

SB2 1 1 12 9 13 13 

SB3 -1 -1 11 8 12 12 

SB4 1 1 11 9 12 12 

SB5 3 3 12 10 13 13 

JE1 2 2 3 3 12 12 

JE2 -12 -12 -5 -5 12 12 

JE3 -6 -6 -5 -5 12 12 

JW1 0 0 4 4 8 8 

JW2 -5 -5 3 3 7 7 

JW3 -5 -5 3 3 7 7 

JW4 -12 -12 3 3 7 7 

PL11 -12 -12 -4 -8 1 2 

SC1 2 2 4 5 10 10 

SC2 4 4 5 5 10 10 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-14 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-4 Risk Analysis Results With 100% Pumping 

Elevations NAVD88 2004.65 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 -1 -1 0 0 6 6 

OW2 -3 -3 3 3 8 8 

NOE1 0 0 2 1 12 13 

NOE2 -5 -5 1 1 11 12 

NOE3 -5 -5 1 1 11 12 

NOE4 -3 -3 2 2 11 12 

NOE5 -11 -11 -2 -2 11 12 

OM1 -12 -12 2 1 12 12 

OM2 -12 -12 0 -12 10 10 

OM3 -12 -12 1 -1 11 11 

OM4 -1 -1 3 -2 13 13 

OM5 -12 -12 1 -2 12 12 

SB1 -12 -12 11 8 12 12 

SB2 1 1 12 9 13 13 

SB3 -3 -3 11 8 12 12 

SB4 1 1 11 9 12 12 

SB5 3 3 12 10 13 13 

JE1 2 2 3 3 12 12 

JE2 -12 -12 -12 -12 11 11 

JE3 -10 -10 -5 -5 12 12 

JW1 0 0 4 4 8 8 

JW2 -5 -5 3 3 7 7 

JW3 -12 -12 3 3 7 7 

JW4 -12 -12 3 3 7 7 

PL11 -12 -12 -5 -12 1 2 

SC1 2 2 4 5 10 10 

SC2 3 3 4 4 9 9 

Human and Economic Risks 

The water elevations and exceedence rate presented in the foregoing sections were combined 

with the stage-damage and stage-fatality relationships developed by the IPET Consequence 

Team to determine the risks to life and property under the scenarios studied. Reference is made 

to the IPET Volume 7 for a complete explanation of the analyses conducted in support of loss 

relationships used herein. 

Economic Risks - The economic consequences considered by the Consequence Team 

included: 1) direct property damages, and 2) indirect economic impacts on local and regional 

economies. Direct property damages were monetary damages to property at risk such as 

residential, commercial, industrial, public buildings, vehicles and infrastructure. For the risk 

scenarios, both pre- and post-Katrina stage-damage functions for properties, are based on base 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-15 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

property conditions that existed prior to Katrina. Only the direct economic damages were used in 

the risk assessment. 

The potential economic damages for the Pre-K and Current HPS without pumping 

(Table 13-5) are equal at about $11.3 Billion for the 50-Year event. The 100-Year potential 

damages are $42.5 Billion for the Pre-K HPS and are reduced by more than $2 Billion to about 

$40.1 Billion for the Current HPS. For the 500-Year event, the sub-basins in both scenarios are 

expected to fill so the potential damages are equal at about $73.3 Billion. 

The potential economic damages for the Pre-K and Current HPS with 50% pumping 

(Table 13-6) are equal at about $3.1 Billion for the 50-Year event. This represents a significant 

$8 Billion reduction from the no pumping scenario and demonstrates the importance of 

maintaining pumping operations during the more frequent events. The 100-Year potential 

damages are $34.1 Billion for the Pre-K HPS and are reduced by almost $10 Billion to about 

$30.3 Billion for the Current HPS, again demonstrating the effectiveness of the pumping system 

at reducing potential damages. For the 500-Year event, the sub-basins in both scenarios are 

expected to fill so the potential damages are equal at about $72.0 Billion. 

The potential economic damages for the Pre-K and Current HPS with 100% pumping 

(Table 13-7) are equal at about $2.4 Billion for the 50-Year event. The 100-Year potential 

damages are $32.5 Billion for the Pre-K HPS and are reduced to about $28.7 Billion for the 

Current HPS. For the 500-Year event, the sub-basins in both scenarios are expected to fill so the 

potential damages are equal at about $72.0 Billion. As may be expected, these values are all less 

than for the 50% pumping scenarios, however, the 100 % pumping scenario is presented for 

comparison purpose only to show how increased efficiency of the pumping system can reduce 

damages. Note that the most significant reductions were attained by the 50% pumping scenario. 

Human Risks - The Consequence Team also examined the human health and safety 

consequences for a variety of event scenarios. For the actual Katrina scenario, the effects 

considered include recorded mortality as well as actual and potential morbidity, including both 

physical and mental health impacts. For the risk scenarios, both pre- and post-Katrina stagefatality functions were developed and provided to the Risk Team for modeling residual risks in 

greater New Orleans. The post-Katrina stage-fatality functions, which are based on population 

and structure conditions that existed prior to Katrina. The mortality risk estimates do NOT 

consider the effects of evacuation on the rate of fatalities within the population at risk. Therefore, 

the fatalities shown in Tables 13-8, 13-9 and 13-10 are potential fatalities only and are not 

intended to be predictions of actual loss of life. 

The potential fatalities for the Pre-K and Current HPS without pumping are equal at about 

492 for the 50-Year event. The 100-Year potential fatalities are 3,708 for the Pre-K HPS and are 

reduced to about 2,388 for the Current HPS. For the 500-Year event, the potential fatalities are 

40,877 for the Pre-K HPS and increase to 42,665 for the Current HPS. The increase in potential 

fatalities for the Current HPS is due to higher levee elevations in some sub-basins, which allow 

more water to collect in the basins. 

The potential fatalities for the Pre-K and Current HPS with 50% pumping are equal at about 

75 for the 50-Year event. This represents a reduction of more than 400 potential fatalities from 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-16 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

the no pumping scenario and demonstrates the importance of maintaining pumping operations 

during the more frequent events. The 100-Year potential fatalities are 2,706 for the Pre-K HPS 

and are reduced to about 1,237 for the Current HPS, again demonstrating the effectiveness of the 

pumping system at reducing potential damages by showing a reduction of 1,000 in potential 

fatalities. For the 500-Year event, the potential fatalities are 29,100 for the Pre-K HPS and 

increase to 30,600 for the Current HPS. The increase in potential fatalities for the Current HPS is 

due to higher levee elevations in some sub-basins, which allow more water to collect in the 

basins. 

The potential fatalities for the Pre-K and Current HPS with 100% pumping are equal at about 

35 for the 50-Year event. The 100-Year potential fatalities are 2,526 for the Pre-K HPS and are 

reduced to about 982 for the Current HPS. For the 500-Year event, the potential fatalities are 

27,700 for the Pre-K HPS and increase to 29,200 for the Current HPS. The increase in potential 

fatalities for the Current HPS is due to higher levee elevations in some sub-basins, which allow 

more water to collect in the basins. As was shown in the economic damages, the potential 

fatalities in the 100% pumping scenario are all less than for the 50% pumping scenarios, but the 

most significant reductions were attained by the 50% pumping scenario. 

Conclusions 

The experience of Katrina proved that the risk to life and property in the New Orleans area 

was high and the results of the risk analysis quantifies the extent of that risk to the Pre-Katrina 

economy and population. The actual direct damages incurred due to the hurricane exceeded 

$28 Billion and the loss of life was more than 700. The corresponding potential damages and life 

loss values obtained by the risk analysis are considerably higher, which points to the severity of 

the risk and attests to the effectiveness of the evacuation prior to the hurricane in reducing the 

loss of life. Examination of the three pumping scenarios shows the importance of the pumping 

system in reducing damages during the more frequent events, but also shows that the system is 

not capable of handling large inflow water volumes from overtopping or breaching during 

extreme events. 

While the HPS has been repaired and improved dramatically over the Pre- Katrina HPS, the 

risk associated with the Current HPS to the area is still considered to be high for extreme events 

if the pre-Katrina potential consequences are used in the analysis. The risks to life and property 

would be expected to be reduced if existing demographics and redevelopment values were used, 

however the reduction would be due entirely to the reduced consequences of system failure and 

not due to the improvements to the system. In any case, the human and economic risks to New 

Orleans would be considered high during exteme events. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-17 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-5 – Potential Damages – No Pumping 

Potential Damages Without Pumping 

In Millions of 2005 Dollars 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 23 23 34 34 170 170 

OW2 412 412 2,305 2,305 3,105 3,105 

NOE1 0 0 8 6 12 12 

NOE2 19 19 123 123 143 143 

NOE3 8 8 510 510 668 671 

NOE4 0 0 42 42 61 62 

NOE5 43 43 4,561 4,561 6,022 6,026 

OM1 402 402 2,209 2,110 2,710 2,710 

OM2 348 348 1,573 962 1,942 1,942 

OM3 376 376 1,709 1,537 2,997 2,997 

OM4 68 68 420 364 1,104 1,104 

OM5 785 785 3,721 2,513 9,596 9,596 

SB1 196 196 2,617 2,511 2,658 2,666 

SB2 - - 25 24 26 26 

SB3 71 71 2,456 2,244 2,496 2,502 

SB4 0 0 497 465 507 511 

SB5 2 2 43 41 44 44 

JE1 420 420 1,093 1,093 5,617 5,617 

JE2 471 471 2,190 2,190 6,270 6,270 

JE3 515 515 4,255 4,255 12,078 12,078 

JW1 0 0 230 230 542 542 

JW2 5 5 393 393 426 426 

JW3 122 122 3,660 3,660 5,625 5,625 

JW4 23 23 5,551 5,551 6,437 6,437 

PL11 33 33 152 90 481 578 

SC1 - - 113 117 132 132 

SC2 390 390 485 485 1,355 1,355 

Totals 4,732 4,732 40,976 38,417 73,224 73,347 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-18 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-6 – Potential Damages – 50% Pumping 

Potential Damages With 50% Pumping 

In Millions of 2005 Dollars 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 23 23 28 28 170 170 

OW2 412 412 2,205 2,205 3,105 3,105 

NOE1 0 0 8 6 12 12 

NOE2 1 1 121 121 143 143 

NOE3 0 0 416 416 667 668 

NOE4 0 0 33 33 60 61 

NOE5 5 5 4,121 4,121 6,007 6,022 

OM1 8 8 2,012 1,864 2,689 2,689 

OM2 - - 1,377 2 1,922 1,922 

OM3 1 1 1,365 1,031 2,953 2,953 

OM4 68 68 420 308 1,096 1,096 

OM5 7 7 2,860 1,483 9,306 9,306 

SB1 - - 2,582 2,403 2,617 2,617 

SB2 - - 25 24 25 25 

SB3 4 4 2,393 2,077 2,456 2,456 

SB4 - - 491 439 497 497 

SB5 2 2 43 41 44 44 

JE1 218 218 420 420 5,122 5,122 

JE2 - - 52 52 6,243 6,243 

JE3 116 116 515 515 12,062 12,062 

JW1 0 0 230 230 542 542 

JW2 1 1 383 383 416 416 

JW3 0 0 2,750 2,750 5,466 5,466 

JW4 - - 5,236 5,236 6,272 6,272 

PL11 - - 0 - 217 245 

SC1 - - 19 19 121 121 

SC2 1,151 1,151 1,355 1,355 1,787 1,787 

Totals 2,017 2,017 31,460 27,563 72,016 72,061 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-19 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-7 – Potential Damages – 50% Pumping 

Potential Damages With 100% Pumping 

In Millions of 2005 Dollars 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 23 23 28 28 170 170 

OW2 412 412 2,205 2,205 3,105 3,105 

NOE1 0 0 8 6 12 12 

NOE2 1 1 121 121 143 143 

NOE3 0 0 416 416 667 668 

NOE4 0 0 33 33 60 61 

NOE5 0 0 4,121 4,121 6,007 6,022 

OM1 - - 2,012 1,864 2,689 2,689 

OM2 - - 1,261 - 1,910 1,910 

OM3 - - 1,031 376 2,910 2,910 

OM4 68 68 420 308 1,096 1,096 

OM5 - - 2,167 200 9,306 9,306 

SB1 - - 2,582 2,403 2,617 2,617 

SB2 - - 25 24 25 25 

SB3 0 0 2,393 2,077 2,456 2,456 

SB4 - - 491 439 497 497 

SB5 2 2 43 41 44 44 

JE1 218 218 420 420 5,122 5,122 

JE2 - - - - 6,216 6,216 

JE3 0 0 515 515 12,062 12,062 

JW1 0 0 230 230 542 542 

JW2 1 1 383 383 416 416 

JW3 - - 2,750 2,750 5,466 5,466 

JW4 - - 5,236 5,236 6,272 6,272 

PL11 - - - - 217 245 

SC1 - - 19 19 121 121 

SC2 1,151 1,151 1,355 1,355 1,787 1,787 

Totals 1,876 1,876 30,266 25,571 71,934 71,979 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-20 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-8 – Potential Fatalities – No Pumping 

Risk Analysis Results Without Pumping 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 0 0 0 0 1 1 

OW2 6 6 76 76 420 420 

NOE1 0 0 0 0 0 0 

NOE2 0 0 1 1 91 154 

NOE3 0 0 13 13 298 524 

NOE4 0 0 0 0 2 2 

NOE5 4 4 63 63 6,528 7,945 

OM1 2 2 80 67 1,923 1,923 

OM2 237 237 1,289 315 3,100 3,100 

OM3 10 10 61 47 1,823 1,823 

OM4 8 8 35 27 484 484 

OM5 21 21 70 49 2,825 2,825 

SB1 - - 915 764 915 915 

SB2 - - - - - - 

SB3 - - 561 455 616 629 

SB4 0 0 82 49 98 102 

SB5 0 0 3 3 3 3 

JE1 0 0 0 0 3 3 

JE2 37 37 50 50 5,367 5,367 

JE3 31 31 73 73 15,271 15,271 

JW1 4 4 31 31 63 63 

JW2 0 0 6 6 40 40 

JW3 1 1 42 42 195 195 

JW4 1 1 132 132 607 607 

PL11 28 28 33 31 180 236 

SC1 - - - - 1 1 

SC2 1 1 1 1 23 23 

Totals 391 391 3,619 2,296 40,877 42,655 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-21 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-9 – Potential Fatalities – 50% Pumping 

Risk Analysis Results with 50% pumping 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 0 0 0 0 1 1 

OW2 6 6 52 52 420 420 

NOE1 0 0 0 0 0 0 

NOE2 0 0 1 1 56 91 

NOE3 - - 9 9 228 298 

NOE4 0 0 0 0 2 2 

NOE5 2 2 45 45 5,111 6,528 

OM1 1 1 54 29 1,241 1,241 

OM2 - - 586 113 2,683 2,683 

OM3 - - 32 24 1,055 1,055 

OM4 8 8 35 20 438 438 

OM5 2 2 56 31 2,075 2,075 

SB1 - - 901 393 915 915 

SB2 - - - - - - 

SB3 - - 561 136 561 561 

SB4 0 0 66 34 82 82 

SB5 0 0 3 3 3 3 

JE1 - - 0 0 3 3 

JE2 - - 23 23 3,716 3,716 

JE3 9 9 31 31 9,678 9,678 

JW1 4 4 31 31 63 63 

JW2 0 0 5 5 24 24 

JW3 0 0 33 33 138 138 

JW4 - - 83 83 418 418 

PL11 0 0 16 1 37 41 

SC1 - - - - - - 

SC2 10 10 23 23 152 152 

Totals 42 42 2,647 1,121 29,099 30,625 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-22 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 13-10 – Potential Fatalities – 100% Pumping 

Risk Analysis Results with 100% pumping 

50-Year elevations 100-Year elevations 500-Year elevations 

Sub-basin Pre-k Current Pre-k Current Pre-k Current 

OW1 0 0 0 0 1 1 

OW2 6 6 52 52 420 420 

NOE1 0 0 0 0 0 0 

NOE2 0 0 1 1 56 91 

NOE3 - - 9 9 228 298 

NOE4 0 0 0 0 2 2 

NOE5 - - 45 45 5,111 6,528 

OM1 - - 54 29 1,241 1,241 

OM2 - - 472 - 2,560 2,560 

OM3 - - 24 10 888 888 

OM4 8 8 35 20 438 438 

OM5 - - 41 12 2,075 2,075 

SB1 - - 901 393 915 915 

SB2 - - - - - - 

SB3 - - 561 136 561 561 

SB4 0 0 66 34 82 82 

SB5 0 0 3 3 3 3 

JE1 - - 0 0 3 3 

JE2 - - - - 2,584 2,584 

JE3 - - 31 31 9,678 9,678 

JW1 4 4 31 31 63 63 

JW2 0 0 5 5 24 24 

JW3 - - 33 33 138 138 

JW4 - - 83 83 418 418 

PL11 0 0 10 0 37 41 

SC1 - - - - - - 

SC2 10 10 23 23 152 152 

Totals 28 28 2,481 951 27,678 29,204 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-13-23 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Inundation Maps 

Figure 13-10. Pre-K Overtopping Rates 
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Figure 13-11 Current Overtopping Rates 
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Figure 13-12 – Jefferson East Bank – Pre-Katrina Inundation Risk – No Pumping 
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Figure 13-13 – Jefferson East Bank – Current Inundation Risk – No Pumping 
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Figure 13-14 – Jefferson East Bank – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure 13-15 – Jefferson East Bank – Current Inundation Risk – 50% Pumping 
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Figure 13-16 – Jefferson East Bank – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure 13-17 – Jefferson East Bank – Current Inundation Risk – 100% Pumping 
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Figure 13-18 – Jefferson West Bank – Pre-Katrina Inundation Risk – No Pumping 
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Figure 13-19 – Jefferson West Bank – Current Inundation Risk- No Pumping 
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Figure 13-20 – Jefferson West Bank – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure 13-21 – Jefferson West Bank – Current Inundation Risk- 50% Pumping 
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Figure 13-22 – Jefferson West Bank – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure 13-23 – Jefferson West Bank – Current Inundation Risk- 100% Pumping 
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Figure 13-24 – New Orleans East – Pre-Katrina Inundation Risk- No Pumping 
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Figure 13-25 – New Orleans East – Current Inundation Risk – No Pumping 
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Figure 13-26 – New Orleans East – Pre-Katrina Inundation Risk- 50% Pumping 
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Figure 13-27 – New Orleans East – Current Inundation Risk – 50% Pumping 
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Figure 13-28 – New Orleans East – Pre-Katrina Inundation Risk- 100% Pumping 
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Figure 13-29 – New Orleans East – Current Inundation Risk – 100% Pumping 
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Figure 13-30 – Orleans East Bank – Pre-Katrina Inundation Risk – No Pumping 
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Figure 13-31 – Orleans East Bank – Current Inundation Risk – No Pumping 
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Figure 13-32 – Orleans East Bank – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure 13-33 – Orleans East Bank – Current Inundation Risk – 50% Pumping 
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Figure 13-34 – Orleans East Bank – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure 13-35 – Orleans East Bank – Current Inundation Risk – 100% Pumping 
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Figure13-36 – Orleans West Bank – Pre-Katrina Inundation Risk – No Pumping 
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Figure13-37 – Orleans West Bank – Current Inundation Risk – No Pumping 
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Figure13-38 – Orleans West Bank – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure13-39 – Orleans West Bank – Current Inundation Risk – 50% Pumping 
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Figure13-40 – Orleans West Bank – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure13-41 – Orleans West Bank – Current Inundation Risk – 100% Pumping 
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Figure 13-42 – Plaquemines (PL 11) – Pre-Katrina Inundation Risk – No Pumping 
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Figure 13-43 – Plaquemines (PL 11)– Current Inundation Risk – No Pumping 
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Figure 13-44 – Plaquemines (PL 11) – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure 13-45 – Plaquemines (PL 11) – Current Inundation Risk – 50% Pumping 
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Figure 13-46 – Plaquemines (PL 11) – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure 13-47 – Plaquemines (PL 11) – Current Inundation Risk – 100% Pumping 
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Figure 13-48 – St. Bernard – Pre-Katrina Inundation Risk – No Pumping 
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Figure 13-49 – St. Bernard – Current Inundation Risk – No Pumping 
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Figure 13-50 – St. Bernard – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure 13-51 – St. Bernard – Current Inundation Risk – 50% Pumping 
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Figure 13-52 – St. Bernard – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure 13-53 – St. Bernard – Current Inundation Risk – 100% Pumping 
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Figure 13-54 – St. Charles – Pre-Katrina Inundation Risk – No Pumping 
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Figure 13-55 – St. Charles – Current Inundation Risk – No Pumping 
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Figure 13-56 – St. Charles – Pre-Katrina Inundation Risk – 50% Pumping 
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Figure 13-57 – St. Charles – Current Inundation Risk – 50% Pumping 
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Figure 13-58 – St. Charles – Pre-Katrina Inundation Risk – 100% Pumping 
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Figure 13-59 – St. Charles – Current Inundation Risk – 100% Pumping 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-1 

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Appendix 14 

Hurricane Protection System 

The Pre-Katrina and Current hurricane protections system (HPS) have been defined in the 

risk analysis using information from USACE design documents, Task Force Guardian, other 

IPET teams and extensive field surveys conducted by the IPET Risk Team. The Pre-Katrina 

analysis is intended to examine the actual risks based on the HPS that existed in New Orleans 

prior to the storm. It is not intended to recreate the “as-designed” HPS and therefore considers 

the failure modes that have been studied by IPET, the actual elevations of walls and levees, the 

in-place engineering properties of the system components and the foundation conditions that 

existed before Katrina struck. 

The current hurricane protections system (HPS) was defined as the system that existed on 

1 June 2007 using information from the New Orleans District, USACE design documents for 

improvements made to the HPS since hurricane Katrina, Task Force Guardian, Task Force Hope 

other IPET teams and the extensive field surveys conducted by the Risk Team. The Pre-Katrina 

HPS risk model was modified to reflect the changes made to the system to repair damage and 

make critical improvements. The revised model is intended to examine the current risks based on 

the HPS that now exists in New Orleans. It considers the failure modes that have been studied by 

IPET, the actual elevations of walls and levees, the in-place engineering properties of the system 

components and the foundation conditions that were found to exist on 1 June 2007. 

Parish Basins and Sub-basins 

The Pre-Katrina and Current HPS limits follow the six main parish boundaries, and each 

parish is subdivided into drainage sub-basins that define how water is collected within the parish. 

Figure 14-1 depicts the parish and sub-basin boundaries that make up the HPS. Boundaries 

between adjacent sub-basins are modeled to allow interflow between them based on the 

elevations of physical features that separate the sub-basins. Sub-basin definitions and interflow 

elevations were provided by the IPET Interior Drainage and Pumping Team. The physical 

features of each parish included in the Pre-Katrina and Current HPS used in the risk model are 

discussed in this appendix, and detailed information for each parish is provided in subsequent 

appendices. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-2 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Reaches 

The boundaries of the basins that provide hurricane protection are divided into reaches that 

have similar engineering properties including structure type, failure modes, materials of 

construction and foundation type. These reaches are shown in tables for each basin that provide: 

• Reach length 

• Reach elevation in NAVD88 2004.65 

• Structure type (W=Wall, L = Levee) 

• Foundation type (H=Hydraulic fill, C=Clay, P=Pile) 

• Sub-basin reference was water from the reach will flow into. 

Three reaches were added to the Current HPS to model the canal closure gates at the end of 

the 17th Street, London Avenue and Orleans Avenue drainage canals. These reaches are numbers 

136, 137 and 138 in the Current HPS. Table 14-18 compares the Pre-Katrina and Current HPS 

crest elevations. All other reach physical characteristics are identical in both HPS. Note 

differences in the fragility of reaches are discussed in Appendix 10. 

Transitions 

HPS components that form part of the flood barrier such as: structural transitions from levee 

to concrete wall, drainage structures, ramps that must be sandbagged during a flood event, 

closure gates and pump stations are included in the risk model. Tables are provided for each 

sub-basin that show: 

• Type of component (T= transition, D=Drainage structure, R=Ramp, G=gate and 

P=Pump station) 

• Reach where the component is located 

• Width of the potential opening 

• Invert elevation for water flowing through the component 

• Sub-basin reference was water from the component will flow into. 

• A description of the component. 

The physical characteristics of transitions are identical in the Pre-Katrina and Current HPS. 

Features 

Additional gate and other features of the HPS are included in the risk model and tables are 

provided that show: 

• Type of component (G= gate and O=Other type of structure) 

• Reach where the component is located 

• Other feature that the component is correlated with. 

• Width of the potential opening 

• Invert elevation for water flowing through the component 

• Sub-basin reference where water from the feature will flow into. 

• A description of the component. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-3 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

The physical characteristics of features are identical in the Pre-Katrina and Current HPS. 

Figure 14-1. Pre-Katrina HPS Basins and Sub-basins 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-4 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

New Orleans East 

The reaches that make up the New Orleans East risk model are shown in Figure 14-2, and 

data are provided in Tables 14-1, 14-2 and 14-3 for reaches, transitions and features, 

respectively. 

Figure 14-2. New Orleans East Reaches 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-5 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-1 New Orleans East Reaches 

Risk Model 

Reach No. 

Reach 

Length (ft) 

Pre-Katrina Elevation 

(NAVD88 2004.65) 

Reach 

Type 

Foundation Material 

Type (H, C, P) 

Subpolder 

Reference (3) 

1 2,405 10.80 W H NOE5 

2 250 10.80 L H NOE5 

3 2,325 10.80 W H NOE5 

4 2,330 10.80 L H NOE5 

5 2,270 10.80 W H NOE5 

6 19,110 13.00 L H NOE5 

7 1,475 13.00 W H NOE5 

8 2,725 13.00 L H NOE5 

9 32900 18.20 L H NOE1 

10 5,830 13.80 L H NOE1 

11 13,325 14.00 L H NOE1 

12 8,910 15.00 L H NOE1 

13 9,185 15.80 L H NOE1 

14 2,615 16.00 L H NOE1 

15 4,470 16.00 L H NOE1 

16 13,045 16.00 L H NOE1 

17 10,570 16.00 L H NOE2 

18 10,760 17.90 W H NOE2 

19 9,320 17.90 W H NOE3 

20 7,905 16.00 L H NOE3 

21 5520 16.00 W H NOE3 

22 385 16.00 L H NOE3 

23 15,320 13.90 L H NOE4 

24 2,910 13.80 W H NOE4 

25 3,230 13.80 L H NOE4 

26 1,640 13.80 W H NOE4 

27 2,750 13.80 L H NOE4 

28 4,100 12.00 L H NOE4 

29 11,185 13.50 W H NOE5 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-6 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-2 New Orleans East Transitions 

Feature 

Number 

Transition 

Type Reach 

Width 

(ft) 

Elevation (ft) (NAVD88 

2004.65) 

SubBasin Description of Feature 

1 R NOE1 25 9 NOE5 Ramp between RR track and bridge 

2 T NOE3 125 5.0 NOE5 LWT NOE 2/3 

3 T NOE3 80 5.0 NOE5 WLT NOE 3/4 

4 T NOE5 155 5.0 NOE5 LWT NOE 4/5 

5 T NOE5 95 5.0 NOE5 WLT NOE 5/6 

6 T NOE7 140 5.0 NOE5 LWT NOE 6/7 

7 T NOE7 130 5.0 NOE5 WLT NOE 7/8 

8 D NOE9 450 16.5 NOE1 Drainage Structure through levee 

9 D NOE9 830 17.5 NOE1 Drainage Structure through levee 

10 D NOE10 65 14.0 NOE1 Drainage Structure through levee 

11 R NOE10 215 8.0 NOE1 I-10 SB ramps 

12 R NOE11 145 7.0 NOE1 I-10 NB ramps 

13 G NOE11 255 6.0 NOE1 Hwy 11 Gate 

14 D NOE11 75 11.0 NOE1 Drainage Structure 

15 D NOE12 55 15.0 NOE1 Drainage Structure 

16 G NOE12 330 15.0 NOE1 Hwy 90 Gate 

17 D NOE14 120 17.0 NOE1 Drainage Structure 

18 G NOE15 95 14.0 NOE1 CSX RR Transition 

19 P NOE17 870 17.3 NOE2 Pump Station #15 

20 T NOE18 135 5.0 NOE2 LWT NOE 17/18 

21 T NOE19 60 5.0 NOE3 WLT NOE 19/20 

22 R NOE20 75 13.0 NOE3 Saturn Blvd Ramp 

23 T NOE21 140 17.0 NOE3 LWT NOE 20/21 

24 T NOE21 25 5.0 NOE3 WLT NOE 21/22 

25 P NOE23 50 5.0 NOE4 Grant Street PS 

26 R NOE23 40 13.0 NOE4 Ramp in levee at steel plant 

27 R NOE23 40 14.0 NOE4 Ramp in levee at steel plant 

28 T NOE24 75 13.0 NOE4 LWT NOE 23/24 - sheet pile 

transition 

29 T NOE24 80 14.0 NOE4 WLT NOE 24/25 

30 T NOE26 75 13.0 NOE4 LWT NOE 25/26 - sheet pile 

transition 

31 T NOE26 60 13.0 NOE4 WLT NOE 26/27 

32 P NOE26 150 13.0 NOE4 Amid PS - PS #20 

33 R NOE27 70 12.0 NOE4 Ramp to Offload Facility 

34 R NOE27 70 9.0 NOE4 Jourdan Rd Ramp 

35 G NOE27 90 5.0 NOE4 RR spur gate closure adjacent to 

Jourdan Road 

36 G NOE28 100 11.0 NOE4 Gate E-4 

37 G NOE28 100 6.0 NOE4 Gate E-5 

38 G NOE28 195 12.0 NOE4 Gate E-6 

39 G NOE28 135 12 NOE4 Gate E-7 

40 R NOE28 35 12 NOE4 Almonaster Ave Ramp from bridge 

41 G NOE 

28 

80 13 NOE4 Gate E-8 

42 G NOE28 90 13 NOE4 Gate E-9 

43 T NOE29 95 13 NOE5 LWT NOE 28/29 

44 R NOE29 50 13 NOE5 Jourdan Rd Ramp at Dwyer Road 

45 G NOE29 30 6 NOE5 RR Gate near Hayne Blvd 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-7 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-3 New Orleans East Features

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 

2004.65) - Sill 

Elevation Subbasin

Description of 

Feature Notes 

1 G NOE1 1 35 1 NOE5 L-13 Swing gate into 

Lakefront airport by 

IHNC (under bridge) 

2 G NOE1 2 22.0 1.75 NOE5 Gate L14 

3 G NOE1 3 63.0 -0.5 NOE5 Gate L15 

4 G NOE7 4 32.0 -1.5 NOE5 Lincoln Beach FW 

Gate 

5 G NOE11 5 30.0 6 NOE1 Hwy 11 Gate 

6 G NOE12 6 80.0 10 NOE1 Hwy 90 Gate 

7 G NOE15 7 20.0 5.7 NOE1 CSX Gate 

8 G NOE18 8 20.0 9.75 NOE2 Michoud Gate 1 Not used 

9 G NOE18 9 20.0 9.75 NOE2 Michoud Gate 2 Not used 

10 G NOE18 10 20.0 9.75 NOE2 Michoud Gate 3 Open 

11 G NOE18 11 20.0 9.75 NOE2 Michoud Gate 4 Open 

12 G NOE18 12 20.0 9.75 NOE2 Michoud Gate 5 Not used 

13 G NOE18 13 20.0 9.75 NOE2 Michoud Gate 6 Open 

14 G NOE18 14 20.0 9.75 NOE2 Michoud Gate 7 Open 

15 G NOE18 15 20.0 9.75 NOE2 Michoud Gate 8 Not used 

16 G NOE18 16 20.0 9.75 NOE2 Michoud Gate 9 Open 

17 G NOE18 17 20.0 9.75 NOE2 Michoud Gate 10 Open 

18 G NOE18 18 20.0 9.75 NOE2 Michoud Gate 11 Open 

19 G NOE18 19 20.0 9.75 NOE2 Michoud Gate 12 Open 

20 G NOE18 20 20.0 9.75 NOE2 Michoud Gate 13 Open 

21 G NOE18 21 20.0 9.75 NOE2 Michoud Gate 14 Open 

22 G NOE18 22 20.0 9.75 NOE2 Michoud Gate 15 Open 

23 G NOE18 23 20.0 9.75 NOE2 Michoud Gate 16 Open 

24 G NOE18 24 20.0 9.75 NOE2 Michoud Gate 17 Open 

25 G NOE18 25 20.0 9.75 NOE2 Michoud Gate 18 Not used 

26 G NOE19 26 20.0 12.8 NOE3 NASA Gate 1 to 

power plant 

Open 

27 G NOE21 27 20.0 12.8 NOE3 NASA Gate 2 to 

power plant 

Open 

28 G NOE21 28 20.5 6.5 NOE3 Gate closure N-1 for 

Bulk Loading Facility 

Open 

29 G NOE27 29 20.0 7.8 NOE4 RR spur gate closure 

adjacent to Jourdan 

Road 

Open 

30 G NOE28 30 20.0 6.5 NOE4 Gate E-4 Closed 

31 G NOE28 31 20.0 6.5 NOE4 Gate E-5 Open 

32 G NOE28 32 17.0 6.5 NOE4 Gate E-6 Closed - RR 

spur 

33 G NOE28 33 20.0 7.2 NOE4 Gate E-7 Closed 

34 G NOE28 34 37.0 6.5 NOE4 Gate E-8 Open RR 

35 G NOE29 35 35.0 6.5 NOE4 Gate E-9 Closed road 

36 G NOE29 36 15.0 7.2 NOE5 Gate E-10 Open to 

docks 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-8 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

37 G NOE29 37 17.0 4.7 NOE5 Gate E-11 Open, RR, 

heavily 

damaged 

38 G NOE29 38 20.0 5.2 NOE5 Gate E-12 Open road 

39 G NOE29 39 17.0 2.2 NOE5 Gate E-13 Open RR 

40 G NOE29 40 30.0 -0.8 NOE5 Gate E-14 Open road 

Hayne Blvd 

41 G NOE29 41 33.0 9.2 NOE5 Gate E-15 RR gate 

42 G NOE29 42 32.0 5.7 NOE5 Gate L-12 Open road 

Jefferson 

The reaches that make up the Jefferson risk model are shown in Figure 14-3, and data are 

provided in Tables 14-4, 14-5 and 14-6 for reaches, transitions and features, respectively. 

Figure 14-3. Jefferson East Bank Reaches 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-9 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-4 Jefferson East Reaches 

Reach 

No. 

Length 

(ft) 

Pre-Katrina Elevation 

(NAVD88 2004.65) 

Reach Type 

(1) 

Foundation Material Type 

(H, C, P) (2) 

Subbasin 

Reference (3) 

30 6,745 12.80 W C JE3 

31 5,915 13.90 W C JE3 

32 4,945 13.90 W C JE3 

33 36,430 14.40 L C JE3 

34 19,925 15.50 L C JE2 

35 12,300 15.50 W H JE2 

36 4,205 25.30 L C OM4 

37 53090 25.40 L C JE1 

38 2,595 9.60 L C JE3 

Table 14-4 Jefferson West Reaches

108 21496 6.5 L H JW1 

109 13947 7.8 L H JW2 

110 24047 7.0 L H JW2 

111 8180 6.5 W C JW2 

112 1730 4.0 L C JW2 

113 320 8.0 W C JW2 

114 1495 9.0 W C JW3 

115 85639 26.3 L C JW1 

116 3060 9.0 W C JW3 

117 11240 8.0 L C JW3 

118 16370 9.8 L C JW3 

119 22135 12.5 L C JW3 

120 6690 12.0 W C JW3 

121 16120 9.0 L C JW3 

122 26700 8.0 L C JW3 

123 9510 5.0 L C JW3 

124 1165 24.8 W C JW3 

125 20710 24.2 L C JW3 

126 40198 22.0 L C OW2 

127 14550 23.2 L C JW4 

128 28337 5.0 L C JW4 

129 44000 8.3 L H JW4 

135 990 24.8 L H JW4 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-10 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

St. Charles 

The reaches that make up the St. Charles risk model are shown in Figure 14-4, and data are 

provided in Tables 14-7, 14-8 and 14-9 for reaches, transitions and features, respectively. 

Figure 14-4. St. Charles Reaches 

Table 14-7 St. Charles Reaches 

Reach 

No. 

Length 

(ft) 

Pre-Katrina Elevation 

(NAVD88 2004.65) 

Reach Type 

(1) 

Foundation Material Type 

(H, C, P) (2) 

Subbasin 

Reference (3) 

40 11710 10.0 L C SC1 

41 23190 10.0 L C SC1 

42 70465 27.9 L C SC2 

43 9280 20.5 L C SC2 

44 3795 20.3 L C SC2 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-11 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-8 St. Charles Transitions 

Feature 

Number 

Transition 

Type Reach 

Width 

(ft) 

Elevation (ft) (NAVD88 

2004.65) 

SubBasin Description of Feature 

62 P SC1 415.0 10.5 SC2 Bayou Trepagneir DS 

63 G SC1 145.0 10.5 SC1 Gate in levee - no road 

64 G SC1 530.0 10.5 SC1 Gate and sheetpile transion at 

Swept Road 

65 R SC1 30.0 10.5 SC1 Road off Airline Hwy 

66 R SC1 25.0 10.5 SC1 Road off Airline Hwy 

67 T SC2 455.0 12.3 SC1 Goodhope DS 

68 D SC2 543.0 11.8 SC1 Cross Bayou DS 

69 D SC2 510.0 11.3 SC1 St. Rose DS 

70 G SC3 28.0 11.8 SC1 Swing Gate at I-310 

71 T SC3 100.0 11.5 SC1 LWT I-310 FW 

72 T SC3 100.0 11.5 SC1 WLT I-310 FW 

73 D SC3 450.0 11.3 SC1 Almedia Canal Drainage 

74 D SC3 450 11.3 SC1 Walker Canal DS 

75 G SC3 450 11.3 SC1 RR Closure Structure 

Table 14-9 St. Charles Features 

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 2004.65) - 

Sill Elevation Subbasin 

Description of 

Feature Notes

Jefferson East 

55 G SC1 55 35 10.5 SC3 Gate in levee - 

no road 

56 G SC3 56 28 6 SC3 Swing Gate at I310 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-12 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Orleans Metro 

The reaches that make up the Orleans Metro risk model are shown in Figure 14-5, and data 

are provided in Tables 14-10, 14-11 and 14-12 for reaches, transitions and features, respectively. 

Figure 14-5. Orleans East Bank Reaches 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-13 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-10 Orleans East Bank Reaches 

Reach 

No. 

Length 

(ft) 

Pre-Katrina Elevation 

(NAVD88 2004.65) 

Reach Type 

(1) 

Foundation Material Type 

(H, C, P) (2) 

Subbasin 

Reference (3) 

45 12,740 15.50 W H OM2 

46 9,280 14.00 W H OM2 

47 3,155 14.00 L H OM2 

48 9,110 14.00 W H OM2 

49 3,610 14.70 L H OM2 

50 12,130 13.50 W H OM1 

51 3,880 13.50 L H OM1 

52 12,765 13.50 W H OM1 

53 3,030 13.50 L H OM1 

54 2,925 12.00 W C OM2 

55 6,310 18.00 L C OM2 

56 9,940 17.00 L C OM1 

57 2,380 16.50 L C OM1 

58 3,220 16.50 W C OM1 

59 7,605 16.50 L C OM1 

60 1,155 14.40 W H OM1 

61 9,095 13.50 W H OM1 

62 9,170 13.80 W H OM3 

63 1,490 13.80 L H OM3 

64 8,390 13.80 W H OM3 

65 875 20.10 W H OM3 

66 1,980 21.50 L H OM3 

67 8,915 22.50 W H OM3 

68 25,450 23.60 W H OM5 

69 10,780 24.30 L H OM5 

70 14,180 24.80 L C OM5 

71 3,350 25.80 L C OM4 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-14 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-11 Orleans East Bank Transitions 

Feature 

Number 

Transition 

Type Reach 

Width 

(ft) 

Elevation (ft) 

(NAVD88 2004.65) 

SubBasin Description of Feature 

Orleans East Bank 

76 P OM1 355 15 OM2 Pump Station #6 - End of 17th Street 

Canal 

77 P OM1 180 15 OM2 Pump Station OP#10 

78 P OM4 170 15 OM2 Pump Station OP#7 FW - End of 

Orleans Canal 

79 U OM6 165 9.0 OM1 Unprotected area adjacent Pump 

Station #3 

80 P OM6 285 14.6 OM1 Pump Station #4 West FW - Middle 

of London Canal 

81 P OM8 179 13.6 OM1 Pump Station #3 FW - End of 

London Canal 

82 P OM8 355 17.3 OM1 Pump Station #4 East FW - Middle of 

London Canal 

83 G OM10 50 13.8 OM2 Gate at W. Roadway Street 

84 T OM10 180 14 OM2 WLT - OM 10/11 

85 G OM11 62 7.55 OM2 Gate 10 - Topaz Dr 

86 R OM11 58 13.8 OM2 NB Ramp at Canal Blvd 

87 R OM11 62 13.8 OM2 SB Ramp at Canal Blvd 

88 T OM12 50 12.05 OM2 Gate 9 - Marconi Dr. 

89 R OM12 160 16 OM2 Ramp 6 - Lakeshore Dr. 

90 R OM12 55 16 OM1 Ramp Lake Terrace Dr. 

91 R OM12 100 16 OM1 Ramp 5 - Lakeshore Dr. 

92 R OM14 150 16.8 OM1 Ramp 4 - Lakeshore Dr. 

93 R OM15 150 16.8 OM1 Ramp 3 - Lakeshore Dr. 

94 R OM15 150 16.3 OM1 Ramp 2 - Franklin Ave - double wide 

ramp 

95 R OM15 80 14.5 OM1 Ramp 1 - Leroy Johnson Drive 

96 G OM15 60 11.8 OM1 Gate 5 - Navy Reserve 

97 G OM15 145 11.55 OM1 Gate 4 - Navy Reserve 

98 T OM15 100 14 OM1 WLT O 16/15 

99 G OM16 28 10.05 OM1 Road Gate 3 

100 G OM16 33 7.3 OM1 RR - Gate 2 

101 R OM16 30 13.25 OM1 Ramp east of France Road near 

Hickey Bridge 

102 R OM17 30 14.75 OM1 France Road Ramp near Chef 

Mentuer Hwy 

103 G OM18 79 6.25 OM3 Road Gate 7W 

104 G OM18 67.5 7.4 OM3 RR Gate 8W 

105 R OM18 40 9 OM3 Ramp to Bridge 

106 G OM18 47 4.75 OM3 RR Gate 9W 

107 R OM18 35 14.75 OM3 France Rd Ramp 

108 G OM19 30 4.25 OM3 Gate 10W - RR open 

109 R OM19 80 14.75 OM3 France Road Ramp 

110 G OM19 75 2.25 OM3 Gate in Levee at Port of NO 

111 G OM19 75 2.25 OM3 Gate in Levee at Port of NO 

112 P OM20 330 6 OM3 Pump Station near Florida Ave 

Bridge 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-15 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

113 T OM23 25 14 OM3 LWT OM 22/23 

Orleans West Bank 

168 R HA1 50 17 OW2 Chalmette - Lower Algiers Ferry 

Ramp 

169 R HA1 50 17 OW2 Chalmette - Lower Algiers Ferry 

Ramp 

170 R HA1 50 17 OW2 Chalmette - Lower Algiers Ferry 

Ramp 

171 R HA1 35 18 OW2 Ramp to Piers 

172 R HA1 125 15 OW2 Ramp to park area at point 

173 T HA1 420 17 OW2 Flood wall and gates around parking 

area 

178 P HA8 360 18.0 OW1 Pump Station #11 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-16 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-12 Orleans Metro Features 

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 

2004.65) - Sill 

Elevation Subbasin

Description of 

Feature Notes 

Orleans East Features 

57 G OM1 57 22 10.6 OM2 Gate E9 - 

Southern RR 

In I-wall 

58 G OM1 58 10 7 OM2 Gate at OP#10 

Pump Station 

In I-wall 

59 G OM1 59 10 7 OM2 Gate north of I-10 In I-wall 

60 G OM1 60 8 7.3 OM2 Gate E4 - 

Veterans Blvd. 

In I-wall 

61 G OM1 61 8 7.3 OM2 Gate E5 - 

Veterans Blvd 

In I-wall 

62 G OM4 62 8.0 7 OM2 Gate at Harrison 

Ave 

In I-wall 

63 G OM4 63 8.0 7 OM2 Gate at Harrison 

Ave 

In I-wall 

64 G OM4 64 8.0 7 OM2 Gate at Filmore 

Ave 

In I-wall 

65 G OM4 65 8.0 7 OM2 Gate at Filmore 

Ave 

In I-wall 

66 G OM6 66 30 10 OM1 West CSX RR 

gate near Pump 

Station #3 

RR tracks 

67 G OM6 67 8 7 OM1 Gate at Filmore 

Ave 

In I-wall 

68 G OM6 68 8 7 OM1 Gate at Filmore 

Ave 

In I-wall 

69 G OM8 69 30 10 OM1 East CSX RR 

gate near Pump 

Station #3 

RR tracks 

70 G OM8 70 8 7 OM1 Gate at Filmore 

Ave 

In I-wall 

71 G OM8 71 8 7 OM1 Gate at Filmore 

Ave 

In I-wall 

72 G OM8 72 8 7 OM1 Gate at Leon C 

Simon Blvd 

In I-wall 

73 G OM10 73 25 8.8 OM2 W. Roadway St With ramp 

74 G OM10 74 25 8.8 OM2 Gate 15 - Into 

Marina Parking 

 

75 G OM10 75 25 8.8 OM2 Gate 14 - Into 

Marina Parking 

 

76 G OM10 76 25 8.8 OM2 Gate 13 - Into 

Marina Parking 

 

77 G OM10 77 30 8.8 OM2 Gate 12 - 

Entrance to 

Marina 

 

78 G OM10 78 60 6.05 OM2 Gate 11 - 

Lakeshore Dr. 

 

79 G OM11 79 62 7.55 OM2 Gate 10 - Topaz 

Dr 

 

80 G OM12 80 50 12.05 OM2 Gate 9 - Marconi 

Dr. 

 

81 G OM12 81 60.0 -5 OM1 Bayou St John 

Floodgate 

 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-17 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 

2004.65) - Sill 

Elevation Subbasin

Description of 

Feature Notes 

82 G OM14 82 25 10 OM1 Gate 3 UNO 

83 G OM14 83 25 10 OM1 Gate 2 UNO 

84 G OM14 84 25 10 OM1 Gate 1 UNO 

85 G OM15 85 22 11.8 OM1 Gate 5 - Navy 

Reserve 

 

86 G OM15 86 34 11.55 OM1 Gate 4 - Navy 

Reserve 

 

87 G OM16 87 28 10.05 OM1 Road Gate 3 

88 G OM16 88 33 7.3 OM1 RR - Gate 2 

89 G OM17 89 30 5 OM1 Gate in France 

Road 

 

90 G OM17 90 20 7.25 OM1 Gate 1W road 

closed 

 

91 G OM17 91 20 7.55 OM1 Gate 2W RR 

open 

 

92 G OM18 92 30 9.5 OM3 Gate 3W access 

open 

 

93 G OM18 93 30 11.25 OM3 Gate 4W access 

open 

 

94 G OM18 94 30 11.25 OM3 Gate 5W access 

open 

 

95 G OM18 95 30 11.25 OM3 Gate 6W access 

open 

 

96 G OM18 96 79 6.25 OM3 Gate 7W road 

closed 

damaged 

97 G OM18 97 67.5 7.4 OM3 Gate 8W RR 

open 

Failed during 

Katrina 

98 G OM18 98 47 4.75 OM3 Gate 9W RR 

open 

damaged 

99 G OM18 99 80 4.25 OM3 Double Gates - 

France Rd 

Parkway 

damaged 

100 G OM19 100 30 4.25 OM3 Gate 10W - RR 

open 

damaged 

101 G OM19 101 25 4.25 OM3 Gate in levee at 

Port of NO 

Failed during 

Katrina 

102 G OM19 102 25 4.25 OM3 Gate in levee at 

Port of NO 

Failed during 

Katrina 

103 G OM20 103 45 2.5 OM3 Gate in pier 

access 

no damage 

104 G OM20 104 45 2.5 OM3 Gate in pier 

access 

no damage 

105 G OM20 105 25 6 OM3 Gate next to 

pump station 

 

106 G OM20 106 30 2.25 OM3 Road closed Florida Avenue 

Bridge Gate 

W20 

107 OM20 107 20 2 OM3 RR closed 

108 G OM20 108 30 2.25 OM3 RR open Gate W21 

109 G OM20 109 40 2.75 OM3 Road open Florida Ave 

Bridge 

110 G OM20 110 30 7.45 OM3 Road open Florida Avenue 

Wharf 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-18 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 

2004.65) - Sill 

Elevation Subbasin

Description of 

Feature Notes 

111 G OM20 111 25 4.45 OM3 RR open Florida Avenue 

Wharf 

112 G OM20 112 30 7.25 OM3 Road open Florida Avenue 

Wharf 

113 G OM20 113 30 8 OM3 Road open 

114 G OM20 114 30 8.25 OM3 Road open 

115 G OM20 115 30 7.5 OM3 Road open 

116 G OM20 116 30 1.75 OM3 Road open Gate W6 

117 G OM20 117 30 2.35 OM3 RR closed Gate W5 

118 G OM20 118 30 5.35 OM3 Road open Gate W4 

119 G OM20 119 30 5.35 OM3 RR Access 

closed 

Gate W3 

120 G OM20 120 30 3.5 OM3 Road open Gate W2 

121 G OM20 121 30 2.25 OM3 Road open Gate W1 

122 G OM23 122 30 7.5 OM3 Road Access Off Poland Ave - 

Navy Complex 

123 G OM23 123 20 7.5 OM3 Road Access Off Poland Ave - 

Navy Complex 

124 G OM23 124 20 7.5 OM3 RR Gate Pauline St 

Wharf 

125 G OM23 125 35 7.5 OM3 Road Access Pauline St 

Wharf 

126 G OM23 126 20 7.5 OM3 RR Gate Pauline St 

Wharf 

127 G OM23 127 20 7.5 OM3 RR Gate Off Charles St. - 

photo 

128 G OM23 128 30 7.5 OM3 Road Access Off Charles St. 

129 G OM23 129 30 7.5 OM3 Road Access Off Charles St. 

130 G OM24 130 20 7.5 OM5 RR Gate Press St. Wharf 

131 G OM24 131 50 7.5 OM5 Road Access Esplande St and 

Wharf 

132 G OM24 132 25 7.5 OM5 Road Access To riverfront 

parking off N 

Peters 

133 G OM24 133 25 7.5 OM5 Road Access To riverfront 

parking - St 

Peters 

134 G OM24 134 25 7.5 OM5 Road Access To riverfront 

parking - 

Toulouse 

135 G OM24 135 25 7.5 OM5 Road Access To riverfront 

parking - St. 

Louis 

136 G OM24 136 25 7.5 OM5 Road Access To riverfront 

parking - Conti 

St 

137 G OM24 137 25 7.5 OM5 Road Access To riverfron 

parking - 

Bienville 

138 G OM24 138 15 7.5 OM5 Pedestrian 

Crossing 

North end of 

Riverwalk 

139 G OM24 139 25 7.5 OM5 Road Access Convention 

Center openings
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-19 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 

2004.65) - Sill 

Elevation Subbasin

Description of 

Feature Notes 

140 G OM24 140 25 5 OM5 Road Access Henderson 

Street 

141 G OM24 141 25 5 OM5 Road Access Race Street 

142 G OM24 142 25 5 OM5 Road Access Orange Street 

143 G OM24 143 50 5 OM5 Road Access Celeste St - 

photo 

144 G OM24 144 50 5 OM5 Road Access Port of NO - 

near Felicity St 

145 G OM24 145 30 5 OM5 Road Access Port of NO - 3rd 

St 

146 G OM24 146 30 5 OM5 Road Access Port of NO - 

Washington St 

147 G OM24 147 30 5 OM5 RR Gate Port of NO - 

across from 9th 

ST 

148 G OM24 148 25 5 OM5 Road Access Port of NO 

149 G OM24 149 20 5 OM5 RR Gate Port of NO 

150 G OM24 150 25 7.5 OM5 Road Access Port of NO - 

Louisana Ave 

151 G OM24 151 25 7.5 OM5 Road Access Port of NO 

152 G OM24 152 20 7.5 OM5 RR Gate Port of NO - 

Napoleon Ave 

153 G OM24 153 30 7.5 OM5 Road Access Port of NO - 

Wharehouse Rd

154 G OM24 154 30 7.5 OM5 Road Access Port of NO 

155 G OM24 155 30 7.5 OM5 Road Access Port of NO - 

Coffee Dr 

156 G OM24 156 30 7.5 OM5 Road Access Port of NO - 

Leake Ave 

157 G OM24 157 30 7.5 OM5 Road Access Port of NO - 

Henry Clay Dr 

Orleans West Bank Features 

191 G HA1 191 30 17 OW2 Gate WB1 - 

Canal Street 

Ferry 

Canal Street 

Ferry 

192 G HA1 192 30 17 OW2 Gate WB2 - 

Canal Street 

Ferry 

Canal Street 

Ferry 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-20 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

St. Bernard 

The reaches that make up the St. Bernard risk model are shown in Figure 14-6, and data are 

provided in Tables 14-13, 14-14 and 14-15 for reaches, transitions and features, respectively. 

Figure 14-6. St. Bernard Reaches 

Table 14-13. St. Bernard Reaches 

Reach 

No. 

Length 

(ft) 

Pre-Katrina Elevation 

(NAVD88 2004.65) 

Reach Type 

(1) 

Foundation Material Type 

(H, C, P) (2) 

Subbasin 

Reference (3) 

72 6,570 13.80 W H SB1 

73 1,115 13.30 W H SB2 

74 26,995 13.60 L H SB2 

75 84,195 15.50 L H SB5 

76 44,650 15.70 L H SB4 

77 25,545 22.00 L H SB4 

78 26,950 21.20 L H SB3 

79 15,885 20.50 L H SB1 

80 870 22.00 W H SB1 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-21 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-14. St. Bernard Transitions 

Feature 

Number 

Transition 

Type Reach 

Width 

(ft) 

Elevation (ft) (NAVD88 

2004.65) 

SubBasin Description of Feature 

114 R SB1 30.0 9 SB1 Ramp for Surekote Rd 

115 P SB2 150.0 14 SB1 Pump Station #5 

116 G SB3 275.0 6 SB2 Gate in levee near Treasure Street 

117 G SB4 255.0 6 SB5 Transition Gate - near Bienvenue 

CS 

118 D SB4 430.0 -3 SB5 Transition Control Structure - Bayou 

Bienvenue 

119 D SB4 455.0 -3 SB5 Transition Control Structure - Bayou 

Dupre 

120 R SB5 170.0 6 SB4 Paris Road (Route 46) ramp 

121 G SB5 155.0 6 SB4 Bayou Road gate transition 

122 P SB5 230.0 15 SB4 Transition Pump Station #8 - St. 

Mary's 

123 R SB5 20.0 6 SB4 Ramp - Dean Road 

124 R SB8 30.0 10 SB4 Ramp at end of River Road 

125 G SB8 195.0 16.5 SB4 Road Gate Closure - canal 

126 G SB8 90 16 SB4 RR Gate Closure - canal 

127 R SB8 30 15 SB4 Ramp for Lumberyard Rd 

128 G SB8 160 16 SB4 RR Gate Closure - canal - 2 RR 

gates 

129 G SB8 40 14.5 SB4 Gate Closure - Domino Sugar Plant 

130 G SB8 45 14 SB4 Gate Closure - Port Ship Service 

Dock 

131 G SB8 65 13.5 SB4 Gate Closure - Near Mehle Ave 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-22 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-15. St. Bernard Features 

Feature 

Number 

Gate (G) or 

Other Point 

Feature (O) Reach 

Correlated 

Features 

Width 

(ft) 

Elevation (ft) 

(NAVD88 2004.65) 

- Sill Elevation Subbasin

Description of 

Feature Notes 

158 G SB1 158 40 6.5 SB1 Gate Closure E-1 damaged 

159 G SB1 159 21 6.5 SB1 RR Gate Closure 

E-2 

damaged 

160 G SB1 160 40 6.5 SB2 Gate Closure S-1 damaged 

161 G SB3 161 25 6.5 SB2 Gate near 

Treasure Street 

no damage

162 G SB4 162 40 6.5 SB2 Gate near Paris 

Road Bridge 

severly 

damaged 

163 G SB4 163 50.0 6.5 SB5 Gate before Bayou 

Bienvenue 

Structure 

damaged 

164 G SB4 164 56.0 -10.8 SB5 Bayou Bienvenue 

Sector Gate 

damaged 

165 G SB4 165 56.0 -10.8 SB5 Bayou Dupre 

Sector Gate 

damaged 

166 G SB5 166 34.0 10.9 SB4 Bayou Road swing 

gate 

damaged 

167 G SB5 167 20 7.4 SB4 RR gate open damaged 

168 G SB5 168 45 6.7 SB4 Highway 39 

closure gate 

damaged 

169 G SB8 169 4 17 SB4 Gate Closure - 

military park 

no damage

170 G SB8 170 32 16.5 SB4 Road Gate 

Closure - canal 

no damage

171 G SB8 171 20 16 SB4 RR gate - canal no damage

172 G SB8 172 24 16 SB4 RR Gate Closure - 

canal 

no damage

173 G SB8 173 24 16 SB4 RR Gate Closure - 

canal 

no damage

174 G SB8 174 20 14.5 SB4 Gate Closure - 

Domino Sugar 

Plant 

no damage

175 G SB8 175 45 14 SB4 Gate Closure - 

Port Ship Service 

Dock 

no damage

176 G SB8 176 35 13.5 SB4 Gate Closure - 

Near Mehle Ave 

no damage
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-23 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Plaquemines 

The reaches that make up the Plaquemines risk model are shown in Figure 14-7a and 

Figure 14-7b, and data are provided in Tables 14-16 and 14-17 for reaches and transitions, 

respectively. There are no point features in the risk model for Plaquemines. 

Figure 14-7a. Plaquemines North Reaches 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-24 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 14-7b. Plaquemines South Reaches 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-25 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-16 - Plaquemines Reaches 

Reach 

No. 

Length 

(ft) 

Pre-Katrina Elevation 

(NAVD88 2004.65) 

Reach Type 

(1) 

Foundation Material Type (H, 

C, P) (2) 

Subbasin 

Reference (3) 

81 22,000 6.00 L H PL11 

82 41,525 8.50 L H PL11 

83 57,470 18.10 L C PL11 

84 50,610 8.50 L H PL2 

85 36,605 16.40 L C PL2 

86 60,615 6.40 L H PL3 

87 25,865 15.70 L C PL3 

88 17,170 11.20 L H PL7 

89 39,195 16.20 L C PL7 

90 27,100 13.50 L H PL7 

91 19,120 13.60 L H PL8 

92 13,774 12.70 L H PL8 

93 6,635 13.80 L H PL8 

94 49,470 16.30 L C PL8 

95 6,160 14.90 L H PL8 

96 26,710 15.00 L H PL9 

97 78,500 14.70 L H PL10 

98 79,100 15.00 L C PL10 

99 22,740 13.90 L C PL9 

100 51,200 16.60 L C PL4 

101 32,235 15.60 L C PL5 

102 50,475 17.30 L H PL4 

103 29,050 17.50 L H PL5 

104 62,810 12.00 L H PL1 

105 30,940 12.40 L H PL6 

106 61,710 18.60 L C PL1 

107 25,225 17.00 L C PL6 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-26 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-17 Plaquemines Transitions 

Feature 

Number 

Transition 

Type Reach 

Width 

(ft) 

Elevation (ft) (NAVD88 

2004.65) 

SubBasin Description of Feature 

132 P PL1 400.0 8 PL12 Pump Station - Belle Chase #2 

133 P PL1 175.0 10 PL12 Pump Station - Belle Chase #1 

134 P PL4 280.0 10 PL2 Pump Station - Upper Ollie 

135 P PL6 100.0 8 PL3 Pump Station - Wilkerson Canal 

136 P PL8 120.0 18 PL8 Pointe A Lache West Pump Station

137 P PL8 170.0 10 PL7 Diamond Pump Station 

138 P PL10 342.0 16 PL11 Hayes Pumping Station 

139 P PL11 550.0 17 PL8 Gainard Woods Pump Station 

140 P PL16 1010.0 20 PL9 Sunrise Pumping Station 

141 G PL16 635.0 19.5 PL9 Empire Flood Gate 

142 P PL17 627.0 19 PL10 Venice Pumping Station 

143 P PL17 975.0 19 PL10 Grand Liard (Buras) PS 

144 P PL22 100 18 PL4 Bellevue Pumping Station 

145 P PL27 175 20 PL5 Pointe A La Hache East Pumping 

Sation 

146 P PL25 100 10.5 PL6 Belair Pump Station 

147 P PL25 200 14 PL1 Scarsdale Pump Station 

148 P PL24 80 9 PL1 Braitwaithe Pump Station 

149 U PL2 2200 8 PL2 Unprotected area between PL11 

and PL2 

150 U PL1 1650 5 PL1 Unprotected area between PLAQ 

and STB 

151 U PL1 1730 18 PL1 MRT between PLAQ and STB 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-27 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 14-18 

Pre-K and Current Reach Elevations 

Levee Heights [ft] 

(1) 

Levee Heights [ft] 

Station (1) 

No. 

Reach 

Name 

Subbasin 

Reference (2) Pre-K Current 

Station 

No. 

Reach 

Name 

Subbasin 

Reference (2) Pre-K Current 

1 NOE 1 NOE5 10.80 10.80 33 JE4 JE3 14.40 16.50 

2 NOE 2 NOE5 10.80 10.80 34 JE5 JE2 15.50 16.50 

3 NOE 3 NOE5 10.80 10.80 35 JE6 JE2 15.50 16.50 

4 NOE 4 NOE5 10.80 10.80 36 JE7 OM4 25.30 26.60 

5 NOE 5 NOE5 10.80 10.80 37 JE8 JE1 25.40 27.00 

6 NOE 6 NOE5 13.00 13.00 38 JE9 JE3 9.60 14.00 

7 NOE 7 NOE5 13.00 13.00 39 SC1 SC1 11.00 13.50 

8 NOE 8 NOE5 13.00 13.00 40 SC2 SC1 10.00 13.00 

9 NOE 9 NOE1 18.20 18.20 41 SC3 SC1 10.00 12.50 

10 NOE 10 NOE1 13.80 13.80 42 SC4 SC2 27.90 28.40 

11 NOE 11 NOE1 14.60 14.00 43 SC5 SC2 20.50 20.60 

12 NOE 12 NOE1 15.00 15.00 44 SC6 SC2 20.30 20.70 

13 NOE 13 NOE1 15.80 15.80 45 OM1 OM2 15.50 16.50 

14 NOE 14 NOE1 17.00 16.00 46 OM2 OM2 14.00 18.00 

15 NOE 15 NOE1 17.30 18.00 47 OM3 OM2 14.00 18.00 

16 NOE 16 NOE1 15.80 18.00 48 OM4 OM2 14.00 18.00 

17 NOE 17 NOE2 17.30 18.00 49 OM5 OM2 14.70 18.00 

18 NOE 18 NOE2 17.90 17.90 50 OM6 OM1 13.50 18.00 

19 NOE 19 NOE3 17.90 17.90 51 OM7 OM1 13.50 18.00 

20 NOE 20 NOE3 17.90 16.00 52 OM8 OM1 13.50 18.00 

21 NOE 21 NOE3 18.50 16.00 53 OM9 OM1 13.50 18.00 

22 NOE 22 NOE3 18.50 16.00 54 OM10 OM2 14.00 14.00 

23 NOE 23 NOE4 13.90 13.90 55 OM11 OM2 18.50 18.50 

24 NOE 24 NOE4 13.80 13.80 56 OM12 OM1 17.00 18.00 

25 NOE 25 NOE4 13.80 13.80 57 OM13 OM1 16.50 18.00 

26 NOE 26 NOE4 13.80 13.80 58 OM14 OM1 16.50 20.50 

27 NOE 27 NOE4 13.80 13.80 59 OM15 OM1 16.50 18.50 

28 NOE 28 NOE4 13.80 13.00 60 OM16 OM1 14.40 18.00 

29 NOE 29 NOE5 13.50 13.50 61 OM17 OM1 13.50 13.50 

30 JE1 JE3 12.80 12.80 62 OM18 OM3 13.80 14.50 

31 JE2 JE3 13.90 13.90 63 OM19 OM3 13.80 14.50 

32 JE3 JE3 13.90 13.90 64 OM20 OM3 13.80 14.50 

 65 OM21 OM3 20.10 22.00 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-14-28 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Levee Heights [ft] 

(1) 

Levee Heights [ft] 

Station (1) 

No. 

Reach 

Name 

Subbasin 

Reference (2) Pre-K Current 

Station 

No. 

Reach 

Name 

Subbasin 

Reference (2) Pre-K Current 

66 OM22 OM3 21.50 21.50 103 PL23 PL5 17.50 17.50 

67 OM23 OM3 22.50 23.70 104 PL24 PL1 12.00 14.50 

68 OM24 OM5 23.60 24.40 105 PL25 PL6 12.40 14.50 

69 OM25 OM5 24.30 25.50 106 PL26 PL1 18.60 18.70 

70 OM26 OM5 24.80 26.00 107 PL27 PL6 17.00 17.60 

71 OM27 OM4 25.80 26.60 108 CW1 JW1 6.5 10.0 

72 SB1 SB1 13.80 15.50 109 CW2 JW2 7.8 10.0 

73 SB2 SB2 13.30 16.50 110 CW3 JW2 7.9 11.0 

74 SB3 SB2 13.60 14.60 111 CW4 JW2 6.5 11.0 

75 SB4 SB5 17.20 18.00 112 CW5 JW2 4.0 11.0 

76 SB5 SB4 15.70 17.20 113 CW6 JW2 8.0 10.0 

77 SB6 SB4 22.00 22.50 114 CW7 JW3 9.0 9.5 

78 SB7 SB3 21.20 21.30 115 CW8 JW1 26.3 27.0 

79 SB8 SB1 20.50 20.50 116 WH1 JW3 9.0 9.5 

80 SB9 SB1 22.00 22.00 117 WH2 JW3 9.5 10.5 

81 PL1 PL1 6.00 10.00 118 WH3 JW3 9.8 11.5 

82 PL2 PL1 8.50 10.00 119 WH4 JW3 12.5 12.5 

83 PL3 PL1 18.10 19.00 120 WH5 JW3 12.5 12.5 

84 PL4 PL2 8.50 8.50 121 WH6 JW3 9.0 10.0 

85 PL5 PL2 16.40 17.50 122 WH7 JW3 9.0 10.0 

86 PL6 PL3 6.40 8.30 123 WH8 JW3 10.0 10.0 

87 PL7 PL3 15.70 17.20 124 WH9 JW3 24.8 24.8 

88 PL8 PL7 11.20 12.80 125 WH10 JW3 24.2 25.0 

89 PL9 PL7 16.20 17.50 126 HA1 OW2 22.0 22.0 

90 PL10 PL7 13.50 13.50 127 HA2 JW4 23.2 24.3 

91 PL11 PL8 13.60 14.00 128 HA3 JW4 9.6 10.0 

92 PL12 PL8 12.70 14.50 129 HA4 JW4 8.3 10.0 

94 PL13 PL8 13.80 15.10 130 HA5 OW2 16.0 21.7 

95 PL14 PL8 16.30 17.50 131 HA6 OW1 20.1 20.8 

96 PL15 PL8 14.90 15.50 132 HA7 OW1 16.0 21.7 

96 PL16 PL9 15.00 15.50 133 HA8 OW1 5.70 10.00 

97 PL17 PL10 14.70 15.50 134 HA9 OW2 8.0 10.0 

98 PL18 PL10 15.00 16.80 135 HA10 JW4 24.80 24.80 

99 PL19 PL9 13.90 17.50 136 CL1 OM2 N/A 16.5 

100 PL20 PL4 16.60 16.70 137 CL2 OM2 N/A 18.0 

101 PL21 PL5 15.60 16.50 138 CL3 OM1 N/A 18.0 

102 PL22 PL4 17.30 17.50 

Notes: (1) Elevation Datum NAVD88-2004.65 : (2) Subbasin where water from overtopping or breaching of reach will collect 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-1 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Appendix 15 

Risk Computer Model -Flood Risk 

Analysis for Tropical Storm 

Environments (FORTE) 

Introduction 

This section describes the Flood Risk Analysis for Tropical Storm Environments (FoRTE) 

tool used to obtain hurricane inundation elevation-exceedence profiles for the New Orleans area. 

FoRTE provides the analytical engine underlying the Interagency Performance Evaluation Task 

Force (IPET) study of the risks associated with the New Orleans hurricane protection system. 

FoRTE was designed to be accessible on most personal computers by leveraging the common 

Microsoft Excel interface. Currently, FoRTE is compatible with Microsoft Excel XP and 2007. 

General Overview and User Interface 

The standard FoRTE user interface is shown in Figure 1 with inputs labeled and described in 

Table 1. In general, execution of the FoRTE tool requires the following three steps: 

1. Input system definition: this step defines the stage-storage relationships for the 

sub-basins, conditions for interflow between adjacent sub-basins, reach, transition, and 

feature definitions, and storm data. 

2. Specify analysis parameters: this step specifies the parameters for analysis, to include 

uncertainty inputs, stratification inputs, and the hydrograph start time. 

3. Specify output options: this step chooses the output and calculation options. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-2 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-1. FoRTE User Interface 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-3 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 15-1. Description of FoRTE Inputs 

Item Description 

A Number of evenly-spaced stratifications of the distribution on surges and waves. The check box to the right of this input 

field turns stratifications on (checked) and off (unchecked). An unchecked box sets the default number of stratifications to 

1 regardless of the value entered in this cell. 

B Log-mean on the uncertainty distribution for surge height. The check box to the right of this input field toggles the 

consideration of uncertainty in surge height, where on (checked) accounts for uncertainty, and off (unchecked) assumes 

no uncertainty. 

C Log-standard deviation on the uncertainty distribution for surge height. This field is ignored if the check box in item B is 

set to off. 

D Log-mean on the uncertainty distribution for wave height. The check box to the right of this input field toggles the 

consideration of uncertainty in wave height, where on (checked) accounts for uncertainty, and off (unchecked) assumes 

no uncertainty. 

E Log-standard deviation on the uncertainty distribution for wave height. This field is ignored if the check box in item D is 

set to off. 

F Prefix for the output file containing surge heights and water volumes for each stratification. The check box to the right of 

this input field determines whether this type of output file will be generated by the FoRTE system (on is checked, and off 

is unchecked). 

G Prefix for the output files containing detailed calculations for each stratification. A separate file is generated for each 

stratification. The check box to the right of this input field determines whether this type of output file will be generated by

the FoRTE system (on is checked, and off is unchecked). 

H Prefix for the output file containing detailed branch output per storm. This file is required for use with the FoRTE storm 

aggregator tool. The check box to the right of this input field determines whether this type of output file will be generated 

by the FoRTE system (on is checked, and off is unchecked). 

I Prefix for the output file containing the aggregate loss exceedance curves for each subbasin based on the number of 

storms studies in a given run. The check box to the right of this input field determines whether results will be aggregated 

to produce loss-exceedance curves, and whether this type of output file will be generated by the FoRTE system (on is 

checked, and off is unchecked). 

J This box turns on storm frequencies. Checked means that frequencies will be used as described in the storm frequencies 

sheet. Unchecked means that the rate is set to one. This latter option is the one needed for aggregating results using the 

FoRTE storm aggregator tool. 

K The starting elevation for generating loss exceedance curves. This input field is ignored if the check box in item I is 

unchecked. 

L The ending elevation for generating loss exceedance curves. This input field is ignored if the check box in item I is 

unchecked. 

M The elevation increment for generating loss exceedance curves. This input field is ignored if the check box in item I is 

unchecked. 

N This is a notes field used to describe the case and system under study. 

O Log standard deviation on the rainfall. This value assumes that rainfall is a lognormally distributed random variable with a 

log mean of 1. 

P Coefficient of variation on the volume of water due to breach for non-overtopping breach failures. This uncertainty is due 

to uncertainty in the Weir coefficient used for calculating water volume. 

Q Coefficient of variation on the volume of water due to overtopping. This uncertainty is due to uncertainty in the Weir 

coefficient used for calculating water volume. 

R Coefficient of variation on the volume of water due to breach for overtopping breach failures. This uncertainty is due to 

uncertainty in the Weir coefficient used for calculating water volume. 

S Coefficient of variation on the volume of water due to open closures and gates. This uncertainty is due to uncertainty in 

the Weir coefficient used for calculating water volume. 

T This is a modification factor used to adjust the height of the hydrographs. This factor is used for epistemic uncertainty 

analysis. The default value of one corresponds to no adjustment of the hydrographs. 

U This is a modification factor used to adjust the position of the fragility curve along the x-axis. This value shifts the entire 

fragility curve along the x-axis. This factor is used for epistemic uncertainty analysis. The default value of zero 

corresponds to no shift in the fragility curve, 

V This is a modification factor used to adjust the value of the Weir coefficients used for calculating volume. This factor is 

used for epistemic uncertainty analysis. The default value of one corresponds to no adjustment to the Weir coefficients. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-4 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

System Definition 

The definition of the hurricane protection system spans several spreadsheets as described in 

the following sections. In particular, the definition of the hurricane protection system includes 

the following elements: 

• High-level basin information that includes the name of the basin and number of 

associated sub-basins; and 

• Stage-storage relationships for each sub-basin that specifies the volume of water held in a 

sub-basin as a function of water elevation; and 

• Interflow mapping matrix that specifies the elevation at which a sub-basin would begin to 

overflow into an adjacent sub-basin; and 

• Reach, transition, and feature data that includes heights, widths, materials, probability of 

gate open for closures, fragility curve for reaches and transitions, and mapping to 

associated reaches (for transitions and closures), sub-basins, and basins. 

Basin Information 

Basic high-level basin information is provided in the “Basin Data” worksheet of the FoRTE 

tool. An annotated snapshot of the “Basin Data” worksheet is provided in Figure 15-2. The 

“Basin Data” worksheet stores the following information: 

• Name of basin 

• Number of sub-basins associated with a basin 

• Prefix for mapping sub-basins and lower-level features to basins 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-5 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-2. Worksheet showing count of sub-basins in each basin. 

Sub-basin Stage-Storage Relationships 

The stage-storage relationships for each sub-basin is provided in the “Sub-basin Data” 

worksheet of the FoRTE tool. An annotated snapshot of the “Sub-basin Data” worksheet is 

provided in Figure 15-3. The “Sub-basin Data” worksheet stores the following information: 

• Vector of water elevations or stage (in feet) for which a corresponding water volume or 

storage is assigned 

• Vector of corresponding water volumes (in cubic feet) for each sub-basin 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-6 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-3. Data input sheet for sub-basin stage-storage relationships. 

Interflow Mapping 

The interflow relationships for each sub-basin is provided in the “Interflow Mapping” 

worksheet of the FoRTE tool. An annotated snapshot of the “Interflow Mapping” worksheet is 

provided in Figure 15-4. The “Interflow Mapping” worksheet stores the following information: 

• Water elevation at which a sub-basin (noted in a row) begins to overflow into an adjacent 

sub-basin (noted in a column). 

Subbasins are across columns

Storage (in cubic feet) 

corresponding to the 

stage in Column A

Stage is in increments 

of 1-foot, spanning a 

range from -30-ft to 

60-ft.
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-4. Sub-basin interflow matrix. 

Reach Definition 

Data that defines the reaches comprising the hurricane protection system is provided in the 

“Reach Data” worksheet of the FoRTE tool. An annotated snapshot of the “Reach Data” 

worksheet is provided in Figure 15-5. Descriptions of the inputs to the “Reach Data” worksheet 

are provided in Table 15-2. 

Interflow relationships are specified in 

the form of a symmetric interflow matrix

Value in these cells 

represent the elevation 

at which the subbasin 

in the row (e.g., 

NOE1) begins to 

overflow into the 

subbasin of the 

column (e.g., NOE2)

Empty cell indicates 

that no interflow 

occurs between the 

row and column 

subbasins. By default, 

diagonal elements are 

blank.
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 15-5. Reach definition worksheet. 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 15-2. Description of Reach Data inputs 

Item Description 

A Reach ID. Each reach is assigned a unique integer ID corresponding to the IDs used to define hydrograph data. 

B Length of the reach section measured in feet. 

C Nominal top elevation of the reach section measured in feet. This is the value used to calculate the volume of water due 

to reach overtopping. 

D Nominal design elevation of the reach section measured in feet. This value is used for specifying failure probabilities on 

the fragility curve. 

E Reach type. “W” corresponds to “Wall” and “L” corresponds to “Levee.” This value is used to determine the appropriate 

Weir coefficient. 

F Reach Weir coefficient. A nominal value of 2.6 is used for levees, and a nominal value of 3.0 is used for walls. 

G This is the ID of the associated basin containing the reach. 

H This is the ID of the associated subbasin containing the reach. 

I Erosion modifier. This value is not currently used for any calculations. 

J Breach fragility curve that specifies the probability of breach as a function of peak water elevation. The low limit 

corresponds to an elevation of 0-feet. The high-limit corresponds to an elevation of 6-feet above the nominal top 

elevation of the reach. Data points specified in between include probability of breach failure at the design and top 

elevations, and 0.5-feet, 1.0-feet, 2.0-feet, and 3.0-feet above the nominal top elevation of the reach. 

K Breach material specifies the composition of the reach as a two-character ID. The first character corresponds to the 

material composition (e.g., “H” for “hydraulic fill”) and the second character corresponds to the length class (e.g., “5” for 

“4000-4999 feet”). This ID is used to determine the breach depth and breach width for use in calculating water volumes 

due to failure. 

L This is the official reach ID as specified by the IPET team. The first set of characters corresponds to the associated 

basin, and the number is a unique ID for reaches in that basin. 

Transition Data 

Data that defines the transitions within the hurricane protection system is provided in the 

“Transition Data” worksheet of the FoRTE tool. An annotated snapshot of the “Transition Data” 

worksheet is provided in Figure 15-6. Descriptions of the inputs to the “Transition Data” 

worksheet are provided in Table 15-3. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 15-6. Transition definition worksheet. 
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This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 15-3. Description of Transition Data inputs 

Item Description 

A Transition ID. Each transition is assigned a unique integer ID. 

B Length of the transition section measured in feet. 

C Nominal top elevation of the transition section measured in feet. This value is used for specifying failure probabilities on 

the fragility curve. 

D Nominal design elevation of the transition section measured in feet. This value is used for specifying failure probabilities 

on the fragility curve. 

E Reach type. “R” corresponds to “Ramp,” “T” corresponds to “Wall-levee,” “D” corresponds to “Drainage Structure,” “P” 

corresponds to “Pumping Stations,” “G” corresponds to “Gates,” and “U” corresponds to “Unknown.” This value is used to 

determine the appropriate breach parameters. 

F Reach weir coefficient. A default value of 2.0 is used for all transitions. 

G This is the IPET ID of the reach containing the transition. This ID is used to map to the appropriate hydrograph. 

H This is the ID of the associated sub-basin containing the transition. 

I This is the FoRTE ID of the reach containing the transition.

J Breach fragility curve that specifies the probability of breach as a function of peak water elevation. The low limit 

corresponds to an elevation of 0-feet. The high-limit corresponds to an elevation of 6-feet above the nominal top 

elevation of the reach. Data points specified in between include probability of breach failure at the design and top 

elevations, and 0.5-feet, 1.0-feet, 2.0-feet, and 3.0-feet above the nominal top elevation of the reach. 

K Transition material is equivalent to reach type in item E above. 

Breach Failure 

Data that define the width and depth of a breach within the hurricane protection system are 

provided in the “Breach Data” worksheet of the FoRTE tool. An annotated snapshot of the 

“Breach Data” worksheet is provided in Figure 15-7. Descriptions of the inputs to the “Breach 

Data” worksheet are provided in Table 15-4. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-12 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-7. Breach data definition worksheet. 

Table 15-4. Description of Breach Data inputs 

Item Description 

A Material and length description. 

B Symbol used for associating different breach materials and lengths to system levees and transitions 

C Breach depths measured from the top of reach or transition (in feet) and breach widths (in feet) for several overtopping 

conditions: (1) 0 to 1-ft overtopping, (2) 1 to 3-ft overtopping, and (3) > 3-ft overtopping. 

D Breach depths measured from the top of reach or transition (in feet) and breach widths (in feet) for non-overtopping 

conditions. Note that these inputs do not apply to transitions. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-13 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Features 

Data that define the closures within the hurricane protection system are provided in the 

“Features” worksheet of the FoRTE tool. An annotated snapshot of the “Features” worksheet is 

provided in Figure 15-8. Descriptions of the inputs to the “Features” worksheet are provided in 

Table 15-5. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-14 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-8. Feature (closure) data definition worksheet. 

A B C D E F G

Each row defines 

a unique gate

H I
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-15 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Table 15-5. Description of Feature Data inputs 

Item Description 

A Feature ID. Each closure is assigned a unique feature ID. 

B Type of feature. Options are “G” for “Gate” and “R” for “Ramp.” 

C Feature category. The only option is “G” for “Gate.” This field is not used for ay calculations. 

D ID of associated reach. This value is used to map the gates to the corresponding reaches. 

E IDs of correlated features used for determining probability of open among a set of related features. 

F Length of closure opening when open (in feet). This value is used with the Wier formula to determine volume of water 

passing through the gate when left open. 

G Bottom elevation of closure when open (in feet). This value is used with the Wier formula to determine volume of water 

passing through the gate when left open. 

H Probability that the gate will be left open during a storm. 

I Associated IPET reach ID. 

Storm Data 

Data that define the storm parameters (not including hydrographs) affecting the hurricane 

protection system are provided in the “Storm Data” worksheet of the FoRTE tool. An annotated 

snapshot of the “Storm Data” worksheet is provided in Figure 15-9. Descriptions of the inputs to 

the “Storm Data” worksheet are provided in Table 15-6. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-16 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-9. Storm data definition worksheet. 

Table 15-6. Description of Storm Data inputs 

Item Description 

A Run ID. This is the ID of the storm. This value is used to map storm parameters to input hydrographs. 

B Storm recurrence rate in events per year. By default this value is set to 1 to accommodate offline aggregation using the 

FoRTE Storm Aggregator. 

C Row ID. This is not a user defined input. 

D Mean volume of water due to precipitation for each storm. This column is repeated for each sub-basin. 

E Standard deviation of water volume due to precipitation for each storm. This value is calculated for each storm and 

sub-basin by multiplying the Rainfall COV by the mean precipitation water volume. 

Hydrograph Processing and Calculation Worksheets 

The FoRTE tool performs calculations on hydrograph data as illustrated in Figure 15-10. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 15-10. Hydrograph processing and calculation worksheets. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-18 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

In particular, FoRTE begins by reading a hydograph file for a given storm into the “Input 

Data” worksheet. Then, for each stratification, FoRTE does the following: 

1. FoRTE applies a stratification factor to the hydrograph surge heights according to the 

current stratification and determines the peak surge for each reach (“Stratified Data” 

worksheet) 

2. The peak surge is determined for each transition (“Transition Surge” worksheet) 

3. The volume of water due to overtopping of each reach is calculated (“Processed Data” 

worksheet) 

4. The volume of water passing through open gates is calculated (“Feature Data” 

worksheet”) 

5. The volume of water due to breach of each reach and transition is calculated (“Breach 

Data” and “Transition Breach Data” worksheets) 

6. The surge and volume data is then accumulated and stored in the “Stratified Inputs” 

worksheet. 

If the option to output “Stratified Water Output per Storm” is selected, the FoRTE tool will 

output the “Stratified Inputs” sheet according to the filename specified on the control sheet user 

interface. 

Branch Calculations and Analysis Results Worksheets 

Following the hydrograph processing and calculation phase, the FoRTE tool processes the 

information for each stratification in turn to determine reach probabilities, and sub-basin water 

volumes, elevations, and probabilities (or rates) for each branch of the system event tree. The 

sheets are described in Figure 15-11. If “Detailed Branch Output per Storm” is desired, the 

FoRTE system will output the “Elevation Consequences” sheet according to the filename 

specified on the control sheet user interface. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Enginee

Figure 15-11. Branch calculations and results worksheets. 

This sheet copies the processed data from the “Stratified Inputs” worksheet according to the cell highlighted above

5

This sheet consolidates all surges and volume calculations for reaches and transitions, and calculates reach-level probabilities including probability of gates being open, probability of breach, and probability of overtopping.

This sheet maps reach data to subbasins, and calculates the water volumes, water elevations, and probability for each branch of the system event tree.

This sheet consolidates the subbasin branch results for each stratification into a single output sheet.

This sheet the elevatioprobabilitieconsidering 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-20 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Performing a FoRTE Analysis 

To perform a FoRTE analysis, perform the following steps: 

1. Enter the appropriate system definition, including sub-basin stage storage and interflow 

relationships, reach data, transition data, breach failure data, and feature data, as was 

described in the previous sections. 

2. Specify analysis parameters and output file options on the control sheet as specified in 

the “General Overview and User Interface” section of this document (Table 15-1). 

3. Click on the “Start Analysis” button. When prompted, select the input hydrographs 

ending with a *.dat extension for calculations. The FoRTE tool accommodates selecting 

as many as 256 data files for batch processing. FoRTE will output files to the same 

directory containing the hydrographs. 

(Note that it may take as long as 1-2 hours per storm depending on the 

capabilities of the host computer; the best performance was observed on a Intel 

Core 2 Duo processor (2 GHz), which resulted in an average run time of about 

20 minutes per storm.) 

4. To consolidate the results from multiple storms to produce loss-exceedance rate curves, 

use the FoRTE Storm Aggregator (Figure 15-12). To do this, proceed as follows: 

a. Load “FoRTE Storm Aggregator” 

b. Check to make sure all storms have the correct frequencies on the “Storm Data” 

worksheet 

c. Click on the “Click Here to Build Loss Exceedance Curves from...” button and select 

the data files corresponding to the storms you wish to aggregate. 

d. When complete, the results will available on the “Elevation Loss Exceedance” 

worksheet. 
Volume VIII Engineering and Operational Risk and Reliability Analysis – Technical Appendix VIII-15-21 

This report is the independent opinion of the IPET and is not necessarily the official position of the U.S. Army Corps of Engineers.

Figure 15-12. Screenshot of the FoRTE Storm Aggregator tool.