Research, Development, and Technology

Turner-Fairbank Highway Research Center

6300 Georgetown Pike

McLean, VA 22101-2296

PUBLICATION NO. FHWA-HRT-24-038 FEBRUARY 2024 

NextScour Case Study: The I–6064/I–95 

Bridge Replacements Over the Lumber River 

in Lumberton, NC 
FOREWORD

The Federal Highway Administration (FHWA) is developing the NextScour Research Initiative 

with the goal of improving scour analysis and providing more accurate scour depth estimates for 

bridge foundation design. This initiative is an interdisciplinary effort between hydraulics and 

geotechnical engineers to address the existing knowledge gaps surrounding traditional methods 

of scour analysis. In 2020, FHWA established a Transportation Pooled Fund study to collaborate 

with State departments of transportation on future bridge projects at sites with layers of cohesive 

soils that could potentially provide resistance to hydraulic loads (FHWA 2020a). This report 

documents one of the case studies, the I–6064/I–95 bridge replacement project in North 

Carolina, and provides an example of how the NextScour approach can assist practitioners as 

they calculate scour depths and conduct bridge foundation design.

Jean A. Nehme, Ph.D., P.E.

Director, Office of Infrastructure

Research and Development

Notice

This document is disseminated under the sponsorship of the U.S. Department of Transportation 

(USDOT) in the interest of information exchange. The U.S. Government assumes no liability for 

the use of the information contained in this document.

The U.S. Government does not endorse products or manufacturers. Trademarks or 

manufacturers’ names appear in this report only because they are considered essential to the 

objective of the document.

Quality Assurance Statement

The Federal Highway Administration (FHWA) provides high-quality information to serve 

Government, industry, and the public in a manner that promotes public understanding. Standards 

and policies are used to ensure and maximize the quality, objectivity, utility, and integrity of its 

information. FHWA periodically reviews quality issues and adjusts its programs and processes to 

ensure continuous quality improvement.
TECHNICAL REPORT DOCUMENTATION PAGE

1. Report No. 

FHWA-HRT-24-038

2. Government Accession No. 3. Recipient’s Catalog No.

4. Title and Subtitle

NextScour Case Study: The I–6064/I–95 Bridge Replacements 

Over the Lumber River in Lumberton, NC

5. Report Date

February 2024 

6. Performing Organization Code:

7. Author(s) with ORCID numbers.

Haoyin Shan (0000-0002-6358-5347), James Pagenkopf (0000-

0001-5392-8628), Jennifer Nicks (0000-0001-7230-3578), 

Zhaoding Xie (0000-0002-4535-2961), Chen Li (0000-0001-

6915-9368), Otto Wiblishauser (0000-0002-1257-8163), Chao 

Huang (0000-0002-6034-8637), and Kornel Kerenyi (0000-

0002-7363-2223).

8. Performing Organization Report No.

9. Performing Organization Name and Address 

Genex Systems, LLC

11848 Rock Landing Dr., Suite 303

Newport News, VA 23606

10. Work Unit No.

11. Contract or Grant No.

693JJ320F000313

12. Sponsoring Agency Name and Address 

Office of Research, Development, and Technology 

Federal Highway Administration

6300 Georgetown Pike

McLean, VA 22101-2296

13. Type of Report and Period Covered

Research Report;

April 2021–November 2022

14. Sponsoring Agency Code 

HRDI-40

15. Supplementary Notes

The Contracting Officer’s Representative was Kornel Kerenyi (HRDI-40; 0000-0002-7363-2223).

16. Abstract

The North Carolina Department of Transportation (NCDOT) plans to replace two highway bridges on I–6064/I–95

over the Lumber River in Lumberton, NC. Hydraulic scour analysis of the planned bridge replacement resulted in 

deep scour depths, which NCDOT believed were excessive due to subsurface layers of cohesive soil that existed 

across the bridge site. NCDOT reached out to the Federal Highway Administration (FHWA) to inquire if FHWA’s 

NextScour Research Initiative could potentially reduce the scour design depths.

Through Transportation Pooled Fund study TPF-5(461), Soil and Erosion Testing Services for Bridge Scour 

Evaluations, FHWA performed four tasks for NCDOT that are summarized in this report (FHWA 2020a). Task 1: 

Perform hydraulic modeling of the site to obtain flow velocities and bed shear stresses. Task 2: Perform soil erosion 

testing on the cohesive soil to determine the clay’s critical shear stress. Task 3: Develop decay functions to establish 

the relation between hydraulic loads and depth. Task 4: Conduct longitudinal scour analysis at the vertical wall at 

the downstream bend. This research study showed that the cohesive layer of clay at the site could potentially stop 

the scour at this bridge site, and the NextScour methodology is a developing tool to improve accuracy for bridge 

foundation design.

17. Key Words

NextScour, scour design, hydraulic modeling, soil 

erosion testing, computational fluid dynamics, and 

probabilistic analysis

18. Distribution Statement

No restrictions. This document is available to the public 

through the National Technical Information Service, 

Springfield, VA 22161.

https://www.ntis.gov

19. Security Classif. (of this report) 

Unclassified

20. Security Classif. (of this page) 

Unclassified

21. No. of Pages

54

22. Price

N/A

Form DOT F 1700.7 (8-72) Reproduction of completed page authorized
ii
iii

TABLE OF CONTENTS

CHAPTER 1. INTRODUCTION................................................................................................ 1

CHAPTER 2. PROJECT BACKGROUND............................................................................... 3

Subsurface Soil Profile and Geotechnical Properties.......................................................... 4

Summary of Current Hydraulic Analysis............................................................................. 5

FHWA 2D Hydraulic Analysis .............................................................................................. 6

CHAPTER 3. TASK 1. HYDRAULIC MODELING................................................................ 7

Nominal Bed Shear Stress Calculation in CFD.................................................................. 11

CHAPTER 4. TASK 2. EROSION TESTING......................................................................... 17

Summary of the ESTD.......................................................................................................... 18

ESTD Erosion Testing Protocol........................................................................................... 19

ESTD Erosion Data Analysis............................................................................................... 20

Critical Shear Stress Distribution ....................................................................................... 25

Supplemental EFA Testing .................................................................................................. 28

NCDOT Additional Soil Investigations............................................................................... 33

CHAPTER 5. TASK 3. DETERMINISTIC SCOUR ANALYSIS......................................... 35

Pier Scour Decay Function Development ........................................................................... 35

Deterministic Pier Scour Analysis....................................................................................... 37

Abutment Scour Decay Function Development................................................................. 38

Deterministic Abutment Scour Analysis............................................................................. 39

Scour Analyses Using HEC-18 Clay Scour Equations ...................................................... 40

CHAPTER 6. TASK 4. LONGITUDINAL SCOUR ANALYSIS AT VERTICAL WALL 

AT DOWNSTREAM BEND...................................................................................................... 43

CHAPTER 7. SUMMARY......................................................................................................... 45

REFERENCES............................................................................................................................ 47
iv

LIST OF FIGURES

Figure 1. Schematic. Profile and plan views of the proposed I–6064/I–95 bridges. ...................... 3

Figure 2. Schematic. Location of the boreholes at the Lumber River. ........................................... 4

Figure 3. Images. Subwatershed elevation (SRH-2D) and computational domain elevation 

(CFD). ................................................................................................................................... 7

Figure 4. Image. Boundary conditions configuration for CFD domain.......................................... 8

Figure 5. Image. Alignment of the four bents................................................................................. 9

Figure 6. Images. Water surface elevation comparison between SRH-2D and CFD models 

for the Q100 flood event. ........................................................................................................ 9

Figure 7. Images. Flow velocity comparison between SRH-2D and CFD models at 

60 percent water depth for the Q100 flood event.................................................................... 9

Figure 8. Images. Bed shear stress comparison between SRH-2D and CFD models for the 

Q100 flood event................................................................................................................... 10

Figure 9. Images. Velocity distribution of each cross section. ..................................................... 11

Figure 10. Images. Shear stress distribution around bents R3 and R4. ........................................ 12

Figure 11. Image. Flow angle of attack and corresponding obstruction width for bents R3 

and R4. ................................................................................................................................ 13

Figure 12. Graph. Cell area used to determine nominal shear stress for bent R4 for the Q500

discharge.............................................................................................................................. 13

Figure 13. Graph. Averaged shear stress over cumulative cell area for bent R4 for the Q500

discharge.............................................................................................................................. 14

Figure 14. Images. Shear stress distribution around abutments for Q100 and Q500 discharge....... 15

Figure 15. Graph. Cell area used to determine the nominal shear stress for the left abutment 

for the Q100 discharge. ......................................................................................................... 15

Figure 16. Graph. Averaged shear stress over cumulative cell area for the left abutment for 

the Q100 discharge................................................................................................................ 16

Figure 17. Schematic. Locations of the boreholes where Shelby tubes were collected. .............. 17

Figure 18. Drawing. NCDOT borings, CPT data, and scour depths. ........................................... 18

Figure 19. Photo. Labeled parts of the ESTD............................................................................... 19

Figure 20. Photo. Laser scan of soil surface and digitized soil surface (inset)............................. 19

Figure 21. Graph. Erosion rates from ESTD testing..................................................................... 21

Figure 22. Graph. Erosion rates from ESTD testing under 50 mm/h. .......................................... 21

Figure 23. Graph. Erosion rates of samples with corresponding erosion categories.................... 22

Figure 24. Photos. ST-11 and ST-23 specimens........................................................................... 23

Figure 25. Photo. ST-3 B-157 specimen....................................................................................... 23

Figure 26. Graph. Erosion rates from ESTD testing of ST-2. ...................................................... 24

Figure 27. Graph. Erosion rates from ST-2 with corresponding erosion categories. ................... 25

Figure 28. Graph. Logarithmic best fit of the erosion data........................................................... 27

Figure 29. Graph. Bootstrapping technique showing 50,000 linear fittings................................. 28

Figure 30. Graph. Histogram of the 50,000 critical shear stresses. .............................................. 28

Figure 31. Photos. ST-15 and ST-29 specimens........................................................................... 29

Figure 32. Sketch. (a) EFA cross section based on Briaud (1999), and (b) laboratory device 

at TFHRC (Nicks et al, 2023). ............................................................................................ 30

Figure 33. Graph. EFA erosion function results, shear stress....................................................... 31

Figure 34. Graph. EFA erosion function results, velocity. ........................................................... 31
v

Figure 35. Graph. Comparison of EFA and ESTD results. .......................................................... 33

Figure 36. Graph. Decay function from FHWA’s flume test data and Annandale and Jones 

data (Annandale 2006). ....................................................................................................... 36

Figure 37. Graph. Deterministic pier scour analysis using decay function and clay 

resistance for Q500................................................................................................................ 38

Figure 38. Graph. Deterministic abutment scour analysis for Q100. ............................................. 40

Figure 39. Image. Shear stress distribution in the Lumber River bend. ....................................... 43

Figure 40. Sketch. Profile sketch for the Lumber River bend. ..................................................... 44
vi

LIST OF TABLES

Table 1. CFD flow condition test matrix. ....................................................................................... 8

Table 2. Flow variables at each cross section for Q100. ................................................................ 11

Table 3. Flow variables at each cross section for Q500. ................................................................ 11

Table 4. Nominal shear stress for hydraulic structures for Q100 and Q500 discharges................... 16

Table 5. Elevations of ESTD erosion test samples and basic soil info......................................... 20

Table 6. Summary of geotechnical results for NCDOT soil specimens. ...................................... 29

Table 7. EFA critical shear stress results...................................................................................... 32

Table 8. Flow variables and scour calculations of ultimate contraction scour and pier scour 

in cohesive soils. ................................................................................................................. 41

Table 9. Shear stress values in the Lumber River bend for various flood events......................... 43
1

CHAPTER 1. INTRODUCTION

The North Carolina Department of Transportation (NCDOT) is investigating the replacement of 

two highway bridges on I–6064/I–95 over the Lumber River in Lumberton, NC. Two hurricanes 

hit this area of North Carolina over the past 6 years, Hurricane Matthew in 2016 and Hurricane 

Florence in 2018. Both occasions resulted in the inundation of I–95 at several locations in 

Lumberton for multiple days (NCDOT 2020). In addition to widening the highway, NCDOT 

plans to raise the low-chord elevation of the proposed bridges. NCDOT calculated scour depths 

for the proposed bridges assuming a uniform layer of sand (i.e., noncohesive material), which 

resulted in significant scour depths. However, they were aware of potential layers of cohesive 

soil material at the bridge site from previous geotechnical investigations. NCDOT reached out to 

the Federal Highway Administration (FHWA) to see whether updated scour depths could be 

predicted by accounting for the cohesive soils at the bridge location and evaluate how the depths 

compared with the traditional scour analysis.

In 2020, FHWA established a Transportation Pooled Fund (TPF) study, TPF-5(461) Soil and 

Erosion Testing Services for Bridge Scour Evaluations (FHWA 2020a). This study is a 

collaborative effort between the J. Sterling Jones Hydraulics Research Laboratory (Hydraulics 

Laboratory) and the Geotechnical Laboratory at Turner-Fairbank Highway Research Center 

(TFHRC) and includes support from the FHWA Resource Center. The study allows State DOTs 

to partner with FHWA to perform the soil and erosion testing required to determine material 

properties, such as critical shear stress, which are needed to calculate scour depths more 

accurately. The I–95 bridges were an excellent candidate for this TPF study. The study also 

aligns with FHWA’s NextScour research initiative, which seeks to improve scour analysis and 

provide more accurate scour depth estimates for foundation design (Shan et al. 2021a).

In addition to soil and erosion testing, FHWA proposed to model the flow around the I–95 bridge

using computational fluid dynamics (CFD) modeling. The Hydraulics Laboratory researchers 

have extensive experience conducting CFD modeling, which captures the complex 

three-dimensional (3D) flow around piers and abutments and provides hydraulic loads (i.e., shear 

stresses) on the riverbed. Additionally, FHWA could help verify the traditional one- and 

two-dimensional (1D and 2D) hydraulic modeling performed for the site.

FHWA outlined the following approach, which contained four tasks:

• Task 1. Hydraulic modeling to obtain the representative initial bed shear stresses near the 

bridge piers and abutments.

• Task 2. Erosion testing to obtain the distribution of clay critical shear stress.

• Task 3. Deterministic scour analysis to obtain decay functions and clay erosion 

resistance.

• Task 4. Scour analysis of the sheet pile retaining wall along the downstream river bend 

adjacent to I–95 northbound (NB).
2

This report references measurements and calculations in both English and International System 

of Units (SI) units. References to bridge plans, hydraulic analysis, and geotechnical studies 

typically use English units. References to lab equipment, lab tests, and CFD modeling typically 

use SI units. Conversions are provided when appropriate.
3

CHAPTER 2. PROJECT BACKGROUND

NCDOT is investigating the replacement of two highway bridges on I–95 (locally I–6064) over 

the Lumber River in Lumberton, NC. The existing NB I–95 bridge was built in 1955 and carries 

two travel lanes, while the existing southbound (SB) bridge was built in 1999 and carries three 

travel lanes. Both bridges have eight spans and are about 380 ft long. The existing bridge surface 

elevation is between 123.2 and 124.8 ft.

Hurricanes Matthew and Florence hit the southeastern region of North Carolina in 2016 and 

2018, respectively, and inundated I–95 at several locations in Lumberton (NCDOT 2020). 

Connectivity of the roadway was disrupted for about 4 days for the Hurricane Matthew event and 

approximately 7 days for Hurricane Florence due to flooding of the Lumber River. Therefore, in 

addition to widening I–95 to eight lanes, NCDOT plans to raise the bridge surface to 135 ft 

through the reach of the interstate spanning Lumberton. The proposed NB and SB bridges are 

about 500 ft long and 177 ft wide in total. Both proposed bridges feature five spans, with a 

120-ft-wide spacing between bents 1 and 2 to accommodate the Lumber River. The low-chord 

elevation of both proposed bridges is about 129 ft. The proposed northeast abutment will be 

placed further back than the current abutment, and the existing soil will be excavated to increase 

the conveyance of flow through the bridge opening. The profile and plan views of the proposed 

bridge are shown in figure 1.

© 2023 NCDOT. Modifications by FHWA.

Figure 1. Schematic. Profile and plan views of the proposed I–6064/I–95 bridges.
4

Each proposed bent features 24 36-inch-diameter cylindrical piles placed at 8.5-ft centers. The 

bents will be positioned at a 60-degree angle from the centerline of the highway to more closely 

align with the thalweg of the Lumber River. The proposed abutments are spill-through stub 

abutments featuring riprap installation.

SUBSURFACE SOIL PROFILE AND GEOTECHNICAL PROPERTIES

NCDOT conducted an initial subsurface investigation of the site in February 2020, consisting of 

borings across the site alternating between the NB and SB lanes. NCDOT made additional 

borings in early 2021. Figure 2 shows borings B-157, B-172, and B-174 from the right and left 

overbanks of the Lumber River. These borings were used to identify the depths to retrieve

undisturbed samples for soil erosion testing.

© 2023 NCDOT. Modifications by FHWA.

Figure 2. Schematic. Location of the boreholes at the Lumber River.

The 2020 borings revealed a layer of alluvial material at the surface, including brown, silty, fine 

sand with trace organic material and gravel. The average elevation in the overbanks was around 

110 ft. The elevation of the channel bed at the thalweg was around 100 ft. Starting at an elevation

around 96 ft was the Coastal Plains layer, also referred to as the Black Creek Formation, a light 

gray, silty clay with traces of mica. This silty clay alternated with layers of light gray, fine to 

coarse sand, with traces of mica down to about 80 ft. Beneath 80 ft was a light gray, fine to 

coarse sand. The layer elevations varied slightly among borings across the site. If the clay layers 

between 96 and 80 ft could potentially provide sufficient erosion resistance to the design floods, 

then the original values of the design scour depths could be reduced.

Standard penetration tests (SPTs) performed in the borings produced blow counts (i.e., N-values) 

of 4–7 blows per foot (bpf) from elevations of 95 to 80 ft, increasing to 12–15 bpf from 75 to 

65 ft. Split-spoon (SS) samples captured material from one of the borings for classification using 

the Unified Soil Classification System (USCS) from ASTM International (ASTM) D2487 

(ASTM 2017). Near B-157, the top 7 ft from the surface (elevation 110 ft) was classified as a 

silty sand (SM) with a median grain size (D50) ranging from 0.15 to 0.22 mm (0.0059 to 0.0087 

inches). Then, starting at elevation 103 ft, the sample was classified as poorly graded sand (SP) 
5

with a D50 of 0.33 mm (0.013 inches). At elevation 96 ft, the SS sample was classified as a 

clayey sand (SC), with a D50 of 0.098 mm (0.0039 inches). At elevation 88 ft, the SS sample was 

classified as high plasticity clay (CH), with a D50 of less than 0.005 mm (0.0002 inches).

SUMMARY OF CURRENT HYDRAULIC ANALYSIS

NCDOT conducted a hydrologic and hydraulic analysis of the site using the 100- and 500-year 

storm events and cross-checked the calculations against the two hurricane events. The Q100

discharge was 14,500 cubic feet per second (cfs), while the Q500 discharge was 20,800 cfs. 

Historical event Hurricane Matthew was considered a 16,500 cfs event, while Hurricane 

Florence was considered a 17,000 cfs event. These flows were verified against U.S. Geological 

Survey (USGS) gage station data (USGS 2023). The total drainage area of the Lumber River 

basin was 725 mi2.

Some unique considerations were noted when constructing the hydraulic model for this site,

including modeling the smaller Five Mile Branch confluence with the Lumber River directly

upstream of the bridge opening. This required splitting the discharge in the hydraulic model into 

two separate inputs. Additionally, a 1.6-mi-long levee was constructed on the downstream 

southeast border of the Lumber River, which had to be modeled correctly to prevent water from

entering that section of the model unless the levee was actually overtopped. Finally, it was noted 

that the bridge at West 5th Street downstream of the reach acted as a constriction that created a 

backwater condition throughout the reach. A USGS gage station located at the bridge provided a 

known outlet condition for the model.

NCDOT modeled the flow parameters of the site using the U.S. Army Corps of Engineers’

(USACE) Hydrologic Engineering Center’s River Analysis System (HEC-RAS) 1D software

(USACE 2022). NCDOT calculated the contraction and pier scour depths using the methodology 

outlined in FHWA’s Hydraulic Engineering Circular No. 18 (HEC-18): Evaluating Scour at 

Bridges (Arneson et al. 2012). The HEC-RAS model returned an average flow depth of 21.17 ft 

at the cross section directly upstream of the bridge for the Q100 flood event, with an average flow 

velocity of 1.65 ft/s. For the Q500 flood, the average flow depth was 23.2 ft and an average flow 

velocity of 1.63 ft/s. NCDOT assumed a uniform sand for the riverbed material, with a D50 grain 

size of 0.2 mm (0.0079 inches), which resulted in critical velocities of 1.61 ft/s for the Q100

discharge, and 1.64 ft/s for the Q500 discharge.

These critical velocities were very close to the actual velocities, which resulted in live-bed scour 

for the 100-year event and clear water scour for the 500-year event. However, per NCDOT 

guidelines, they do not consider the clear water case for nonpressure flow conditions, so only 

live-bed scour equations were considered for both flood events. The predicted live-bed 

contraction scour depths for the Q100 and Q500 discharges were 29.1 ft and 38.3 ft, respectively.

For the Q100 and Q500 discharges, the clear water contraction scour depth calculation was 40.0 ft

and 50.5 ft, respectively.

Local pier scour was calculated for each of the four bents for both the 100- and 500-year events.

For the two bents closest to the main channel, the resulting total pier scour was around 7.5 ft for 

the Q100 discharge. For the two bents in the left floodplain, the pier scour was around 11.4 ft, due 

to higher velocities compared to the main channel. For the Q500 discharge, the local pier scour 
6

depths increased to 8 ft for the bents adjacent to the channel and 12.1 ft for the bents in the 

floodplain. Considering potential channel migration, the deepest local pier scour depths for each 

event were used to calculate total scour.

The total scour depth (including pier and contraction scour) was 40.5 ft for the Q100 discharge

and 50.8 ft for the Q500 discharge. These scour depths were deep enough that the cohesive soil in 

the potential clay layer between 96 and 80 ft could provide a significant reduction in scour depth 

if the soil was capable of providing erosion resistance to the hydraulic loads.

FHWA 2D HYDRAULIC ANALYSIS

FHWA conducted a separate 2D hydraulic analysis of the bridge site using Sedimentation and 

River Hydraulics (SRH-2D) modeling of a Q100 flood event (U.S. Bureau of Reclamation 

(USBR) 2022). The model determined that the average flow depth and average velocity upstream 

of the contraction was 21.73 ft and 1.08 ft/s, respectively. The bed material was still assumed to 

be uniform sand with a D50 of 0.2 mm (0.0079 inches). The model considered main channel 

hydraulics and determined that the width and discharge of the main channel were 100 ft and 

6,128 cfs, respectively. With a low upstream velocity of only 1.08 ft/s, the FHWA Hydraulic 

Toolbox predicted a clear water scour condition due to a critical velocity of 1.62 ft/s (FHWA 

2020b). The results of the live-bed method predicted a contraction scour depth of 26.7 ft, which 

was similar to the live-bed calculation of 29.1 ft for the Q100 event in the previous section. 

However, the results of the clear water method predicted a contraction scour depth of only 9.7 ft, 

which was much less than the depth of 40.0 ft calculated in the previous section.

Per HEC-18, the clear water condition governs due to the low velocity of 1.08 ft/s compared to 

the critical velocity of 1.62 ft/s (Arneson et al. 2012). Using the clear water contraction scour 

depth of 9.7 ft, the total scour depth (including 11.4 ft of pier scour) equaled 21.1 ft for the Q100

discharge. This total scour depth, although shallower than the 40.5 ft calculated by NCDOT, still 

placed it within the range of the potential clay layer between 96 and 80 ft.
7

CHAPTER 3. TASK 1. HYDRAULIC MODELING

FHWA modelers used a 3D CFD simulation to analyze the hydraulic parameters around the piers 

and abutments of the proposed bridge. The simulation aimed to compare the flow characteristics 

and bed shear stresses with the results from the 2D simulation using SRH-2D modeling and to 

verify the design of the proposed hydraulic structures using various flood events (USBR 2022). 

The full-scale 3D CFD simulation requires fewer assumptions than traditional 1D and 2D 

modeling, so the results should be more accurate, but the computational power required to run 

the simulation limits the size of the domain.

The full-scale simulation for NCDOT I–6064/I–95 Lumber River bridge was divided into three 

stages: preparation of the full-scale computational domain, preprocessing of the simulation, and 

postprocessing analysis of the results. In the first stage, the computational domain for the 

Lumber River bridge subwatershed was generated using point cloud data exported from the 

SRH-2D model, as shown in figure 3. NCDOT developed this digital elevation model (DEM) 

using light detection and ranging (LiDAR) data combined with cross-section survey data. Figure 

3 shows the whole subwatershed domain used for SRH-2D modeling, and a polygonal box 

shows the boundary of the smaller CFD domain. Both the SRH-2D and CFD domains included 

the proposed design of the I–95 highway, bents, and spill-through abutments. The CFD domain 

in figure 3 is 3,772 ft long by 2,415 ft wide.

Source: FHWA. DEM data provided by NCDOT.

Figure 3. Images. Subwatershed elevation (SRH-2D) and computational domain elevation 

(CFD).

The second stage of model preparation included setting the boundary condition configuration and 

mesh scheme. Figure 4 illustrates the boundary configurations, where sections A and B represent 

the Lumber River and Five Mile Branch inputs upstream, respectively, as velocity inlets. The 

model outlet is represented by section C as a static pressure outlet with known flow depths. The 

flow conditions at each boundary in the CFD domain were set by the corresponding boundary

conditions from the SRH-2D model. Multiple flood events were considered, which are 
8

summarized in the test matrix in table 1. The volume of fraction (VOF) model was used for the 

CFD simulations, which computes the water surface elevation (WSE) based on the iso value of 

water volume fraction.

Source: FHWA.

Figure 4. Image. Boundary conditions configuration for CFD domain.

Table 1. CFD flow condition test matrix.

USGS Gage 02134170* at 

Lumber River (flood)

Section A Section B Section C

Q (ft3/s) WSE (ft) Q (ft3/s) WSE (ft) WSE (ft)

Q2 2,780 116.27 608 115.83 114.94

Q10 6,080 119.85 1,330 119.7 119.06

Q25 7,900 121.68 2,100 121.55 120.86

Q50 9,350 123.04 2,850 122.92 122.18

Q100 10,900 124.3 3,600 124.18 123.41

Q500 14,700 126.64 6,100 126.54 125.6

*USGS 2023.

The third stage was postprocessing of the simulation, which included analyzing water elevation 

and velocities across the entire domain, bed shear stress around the bents and abutments, and 

velocities across five predetermined cross sections. Figure 5 shows the numbering system for the 

bents from R1 to R4. Each bent consists of a pier cap and 24 cylindrical columns. Figure 6 to 

figure 8 compare the CFD results with the SRH-2D model for the Q100 flood event, including 

water surface elevation, velocity distribution, and bed shear stresses. Figure 6 shows that the

water surface elevation matches closely between the two models. Figure 7 shows higher 

velocities along each bent in the CFD model and slightly higher velocities overall. The biggest 

difference between the two models is bed shear stress. Figure 8 shows that the SRH-2D model 

predicts much higher shear stress values, both within the main channel and especially in the 

overbank areas. The higher shear stress values are due to parameters such as Manning’s n value

and because the 2D model calculates the depth-averaged velocity. The bed shear stresses 
9

computed by the CFD model focus on the near-bed region (i.e., the hydraulic loads that account 

for sediment transport at the site).

Source: FHWA.

Figure 5. Image. Alignment of the four bents.

Source: FHWA.

Figure 6. Images. Water surface elevation comparison between SRH-2D and CFD models 

for the Q100 flood event.

Source: FHWA.

Figure 7. Images. Flow velocity comparison between SRH-2D and CFD models at 

60 percent water depth for the Q100 flood event.
10

Source: FHWA.

Note: 1 Pa = 0.021 pounds per square ft (psf).

Figure 8. Images. Bed shear stress comparison between SRH-2D and CFD models for the 

Q100 flood event.

Figure 9 shows velocity contour plots at five cross sections in the domain, which are labeled P1 

to P5. P1 is positioned at the main channel of the Lumber River, upstream of the bridge, while 

P2 is positioned at the main channel of Five Mile Branch. P3 is situated after the confluence 

directly upstream of the bridge and stretches across the bridge opening. P4 is similar to P3 but 

located at the downstream end of the SB bridge. P5 is positioned in the main channel of the 

Lumber River downstream of the bridge. Average velocities and water surface elevations for 

each cross section are provided in table 2 and table 3 for Q100 and Q500 discharges, respectively. 

The tables also compare results from the CFD and SRH-2D simulations at each cross section.

Table 2 and table 3 show that the average velocity of 2.2 ft/s in the main channel of the Lumber 

River upstream of the bridge was higher than the average velocity in the main channel of Five 

Mile Branch for both flood events. Both of these velocities were smaller than at cross sections P3 

and P4, where the flow discharges were combined. The highest velocity of 3.5 ft/s (Q100) or 4 ft/s 

(Q500) occurred at cross section P5, where the flow was accelerated along the main channel of the 

Lumber River. CFD’s results of water surface elevation for each section were close to 

SRH-2D’s, with an average difference of 0.5 ft. The difference in velocities ranged from 0.3 ft/s 

to 0.5 ft/s except at P5, which had a slightly larger difference of 0.9 ft/s because SRH-2D cannot 

fully model the 3D flow field.
11

Source: FHWA.

Figure 9. Images. Velocity distribution of each cross section.

Table 2. Flow variables at each cross section for Q100.

Cross 

Section

Average Velocity (ft/s) WSE (ft)

CFD SRH-2D CFD SRH-2D

P1 2.2 1.9 123.7 124.1

P2 1.7 1.2 123.6 124

P3 2.8 2.1 123.5 124

P4 2.9 2.4 123.5 123.9

P5 3.5 2.6 123.4 123.8

Table 3. Flow variables at each cross section for Q500.

Cross 

Section

Average Velocity (ft/s) WSE (ft)

CFD SRH-2D CFD SRH-2D

P1 2.2 2.3 126.1 126.3

P2 2 1.5 125.8 126.3

P3 2.9 2.6 125.8 126.2

P4 3 2.9 125.7 126.1

P5 4 3.2 125.6 126

NOMINAL BED SHEAR STRESS CALCULATION IN CFD

The bed shear stress distribution was used to identify potential areas for scour around piers and 

abutments. Based on the observation of shear stresses in figure 8, the areas surrounding bents R3 

and R4 had higher bed shear stresses than the R1 and R2 areas; therefore, the shear analysis 

focused on the R3 and R4 bents, as shown in figure 10.
12

Source: FHWA.

Figure 10. Images. Shear stress distribution around bents R3 and R4.

The methodology used to obtain the nominal shear stress for each bent included three steps. 

First, the shear stress value for each computation cell located within 2 column diameters of bents

R3 and R4 was collected and sorted in descending order. Next, the area-averaged shear stress 

(defined in equation 1) was applied to each cell to obtain a series of averaged shear stresses 

based on the accumulated cell area. Finally, the projected area was used to establish the total cell 

area summation cutoff value, where the projected area was equal to the obstruction width times 

the flow depth. At this site, there was a significant angle of attack from the flow at bents R3 and 

R4 as the flow came around the left abutment. The angles of attack for each bent were 

determined by studying the flow’s vector field from the 2D model. Per HEC-18, a maximum of 

12 cylindrical columns should be considered when calculating the correction factor for attack 

angle of flow, so that value was used in this analysis (Arneson et al. 2012). Lines perpendicular 

to the flow field vectors were estimated starting at each lead column and extending to a right 

angle of the 12th column. This geometry produced angles of attack of 32 and 39 degrees for R3 

and R4, respectively. Not considering the spacing between columns, this resulted in projected 

frontal widths of 19.1 ft for R3 and 22.7 ft for R4, which are shown in figure 11.

The mean shear stress value is given in equation 1.

(1)

Where:

= averaged shear stress for the corresponding number of cells (Pa).

i = the cell number.

i = shear stress for cell i (Pa).

Ai = area of ith cell (m2).

n = cell number where the sum of all cells is equal to projected area.
13

Source: FHWA.

Figure 11. Image. Flow angle of attack and corresponding obstruction width for bents R3 

and R4.

Based on the approach described previously, figure 12 shows the cells selected to compute the 

nominal shear stress at bent R4 for the Q500 discharge. The combined area enclosed within the 

dashed lines is equal to the projected area for bent R4, which was 32.8 m2(352.9 ft2). The

corresponding nominal shear stress value, shown in figure 13, was 5.6 Pa (0.118 pounds per 

square ft (psf)).

Source: FHWA.

Figure 12. Graph. Cell area used to determine nominal shear stress for bent R4 for the 

Q500 discharge.
14

Source: FHWA.

Figure 13. Graph. Averaged shear stress over cumulative cell area for bent R4 for the 

Q500 discharge.

The same approach was adopted for the abutments, as shown in figure 14. The computational 

cells considered were located on the embankment slope from the water surface elevation down to 

the toe of the abutment. For abutments, the projected area was measured as the flow depth times 

the length of the abutment wall to the toe of the abutment slope plus an additional two times the 

flow depth. Figure 15 shows the cells selected to compute the nominal shear stress at the left 

abutment for the Q100 discharge. The averaged shear stress values for these cells are plotted in 

figure 16. The nominal shear stresses at the left and right abutments for the Q100 discharge was 

17.8 Pa (0.374 psf) and 14.9 Pa (0.313 psf), respectively. Table 4 summarizes the values of 

nominal shear stresses and projected areas from the two floods. The maximum nominal shear 

stress value for pier scour is 5.6 Pa (0.118 psf) and occurs at bent R4 during the Q500 flood. The 

maximum nominal shear stress value for abutment scour is 17.8 Pa (0.374 psf) and occurs at the 

left abutment during the Q100 flood.
15

Source: FHWA.

Figure 14. Images. Shear stress distribution around abutments for Q100 and Q500 discharge.

Source: FHWA.

Figure 15. Graph. Cell area used to determine the nominal shear stress for the left 

abutment for the Q100 discharge.
16

Source: FHWA.

Figure 16. Graph. Averaged shear stress over cumulative cell area for the left abutment for 

the Q100 discharge.

Table 4. Nominal shear stress for hydraulic structures for Q100 and Q500 discharges.

Hydraulic 

Structures

Q100 Q500

Projected 

Area (ft2)

Nominal 

Shear Stress 

(Pa)

Projected 

Area (ft2)

Nominal 

Shear Stress 

(Pa)

Bent R3 242.6 4.6 297.6 5.0

Bent R4 288.3 5.0 353.7 5.6

Left abutment 515.3 17.8 215.8 16.5

Right abutment 517.5 14.9 215.8 16.0

In addition to the nominal shear stress values calculated for bents R3 and R4 and the two 

abutments, approach shear stress values were determined from the CFD model for both flood 

discharges. These upstream values were determined by averaging the cells from the P1 cross 

sections across the main channel. The approach shear stress was 2.86 Pa (0.060 psf) for the 

Q100 discharge and 2.52 Pa (0.053 psf) for the Q500 discharge.
17

CHAPTER 4. TASK 2. EROSION TESTING

NCDOT returned to the site in May 2021 to collect Shelby tube soil samples from boreholes 

B-157, B-172, and B-174. The drillers collected 23 Shelby tube soil samples in total, including 5 

from B-157 from elevation 85 to 76 ft, 11 from B-172 from elevation 96 to 74 ft, and 7 from 

B-174 from elevation 98 to 83 ft. The samples were then shipped to FHWA to perform erosion 

testing with the Ex-situ Scour Testing Device (ESTD). FHWA also conducted similar erosion 

testing experiments with the Erosion Function Apparatus (EFA) device in the Geotechnical 

Laboratory. This additional testing would help validate the results from both devices. 

Additionally, basic index tests were conducted in the Geotechnical Laboratory to classify the 

soils tested in the ESTD and EFA. Figure 17 illustrates the locations of the three boreholes on the 

overbanks on either side of the Lumber River. Per FHWA’s suggestion, NCDOT also conducted 

cone penetration tests (CPTs) adjacent to the three boreholes, which are also marked on figure 

17.

© 2023 NCDOT. Modifications by FHWA.

Note: Circles represent boreholes; triangles represent CPT locations.

Figure 17. Schematic. Locations of the boreholes where Shelby tubes were collected.

FHWA researchers selected the samples closest to the surface to conduct erosion testing in the 

ESTD. These Shelby tube (ST) samples included ST-2 from B-157; ST-9, ST-10, and ST-11 

from B-172; and ST-22, ST-23, and ST-24 from B-174. To compare the results of the ESTD to 

the EFA, they selected ST-3 from B-157 to compare to ST-2. They then also selected samples 

ST-15 from B-172 and ST-29 from B-174 to conduct additional erosion testing in the EFA 

because the soil layer appeared similar to ST-2 based on the CPT data. Some discrepancies 

existed between the original bore logs and the CPT data, where the original bore logs showed a 

silty clay layer starting at elevation 96 ft for borings B-172 and B-174, but the CPT data showed 

silty sand mixtures at that depth and silt and clay mixtures beginning at elevation 83 ft. Figure 18

summarizes the CPT data and the soil samples tested in the erosion devices, as well as the river 

elevation and Q100 scour depth.
18

Source: FHWA. CPT data courtesy of NCDOT.

Figure 18. Drawing. NCDOT borings, CPT data, and scour depths.

SUMMARY OF THE ESTD

The ESTD, shown in figure 19, is an automated erosion device developed by FHWA that 

measures the erodibility of a cylindrical soil sample under well-controlled flow conditions (Shan 

et al. 2011). The ESTD features a 4.7-inch-wide by 0.75-inch-high rectangular acrylic test 

channel that is 3 ft long. The maximum flow capacity of the pump is 0.5 ft3/s, which translates to 

a maximum flow velocity of 20 ft/s. An underwater laser scanner mounted on an industrial 

robotic arm scans the soil surface every 20 s (figure 20), sending a quasi-instantaneous signal to 

the control program. The control program averages the scan data and compares it to a reference 

point on the surface of the test channel. If the average value is less than this reference point, a 

command is sent to the piston to extrude the sample to maintain a surface flush with the test 

channel bed. The control program records the position data of the piston, which is used to 

calculate the erosion rate. Flow circulation is continuously measured using an electromagnetic 

flow meter, and two filtration tanks are used to capture the eroded clay particles and keep the 

water clear for the laser scanner. An electromagnetic shear sensor located just upstream of the 

piston directly measures the shear stress of the soil sample in a separate series of tests after 

erosion testing is completed. The relationship between erosion rates and shear stresses was

developed by comparing results at similar flow rates, which are then used to determine the 

critical shear stress of the soil. Shan et al. (2021b) outlined the ESTD test procedures in detail.
19

Source: FHWA.

A = robotic arm; B = laser scanner; C = flow channel; D = 3-inch Shelby 

tube sample; E = hydraulic piston; F = shear sensor; G = flow pump; 

H = flow meter; and I = filter cylinders.

Figure 19. Photo. Labeled parts of the ESTD.

Source: FHWA.

Figure 20. Photo. Laser scan of soil surface and digitized soil surface (inset).

ESTD EROSION TESTING PROTOCOL

For ESTD erosion tests, soil samples are tested in 1-ft sections, typically taken from the bottom 

of the Shelby tube. For the NCDOT Shelby tubes, only one sample was tested from each tube.

Simple strength tests were performed on the top and bottom surfaces of each 1-ft sample, 

including pocket penetrometer (PP) and shear vane tests. The top surface is then extruded, 

trimmed, and mounted in the ESTD in preparation for the erosion tests.

Table 5 provides the sample elevations, water content, and strength properties for the seven 1-ft 

soil samples tested in the ESTD. Additional soil classifications were performed on several of the 

samples by the Geotechnical Laboratory and are presented later in the report.
20

Table 5. Elevations of ESTD erosion test samples and basic soil info.

Sample ID

Elevation 

(ft)

Elevation of 

1-ft Test 

Tube (ft)

Elevation

of Strength 

Test (ft)

qu

*

(kPa)

Su

**

(kPa)

Water 

Content 

(percent)

B-157 ST-2 85.4–83.8 84.8–83.8 — 150, 275 27, 144 15, 25

B-172 ST-9 91.4–89.5 91.4–90.4 — — — 22

B-172 ST-10 89.5–87.5 88.5–87.5 88.5 — — 20

B-172 ST-11 87.5–85.5 86.5–85.5 86.5 450 88 20

B-174 ST-22 98.5–96.5 97.5–96.5 97.5 360 150 17

B-174 ST-23 96.5–94.5 96.5–95.5 — 450 50, 120 14.9, 17.7

B-174 ST-24 94.5–92.5 93.5–92.5 93.5 450 144 15

*Unconfined compressive strength as measured with a PP.

**Undrained shear strength as measured with a vane shear device.

1 kPa = 21 psf.

—No data available.

Erosion tests on each soil sample typically start at 2 L/s (0.071 cfs) and increase in flow rate to a 

maximum of 13 L/s (0.46 cfs) in steps of 1 L/s (0.035 cfs). Each flow rate is tested for 10 min 

and then is automatically increased to the next flow rate by the control program. The initial flow 

rate can be increased if the soil is known to be resistant at low flow rates or to rerun a specific set 

of flow rates. The test ends when the maximum flow rate is reached or the sample runs out of 

material. Typically, only the top 9–10 inches of the soil are tested, as the bottom 2–3 inches can 

be unstable and dislodge during testing. A linear best fit line was applied to the erosion data 

(recorded position of the soil surface versus time) for each flow rate, where the slope represented 

the erosion rate.

Hydraulic shear stresses on soil samples were recorded separately from the erosion tests using 

0.6-inch-thick soil samples extruded into stainless steel rings. The ring was mounted to a small 

bowl that attaches directly to the sensor disk. The bowl’s interior features a raised circular 

platform, which pushes the soil out about 0.08 inches above the edge of the ring. Once mounted 

on the shear sensor, the sensor disk height was adjusted until the soil surface was flush with the 

test channel surface. The shear stress data were used to convert flow rates into equivalent shear 

stress values.

ESTD EROSION DATA ANALYSIS

Generally, the soils within all seven Shelby tube samples had higher erosion rates as the flow 

rates increased (i.e., increased shear stresses), as shown in figure 21. A few of the samples, 

including ST-10 and ST-11 from B-172 and ST-23 and ST-24 from B-174 had a wider range of 

erosion rates than the other samples. Results from index tests and soil classifications performed 

by the Geotechnical Laboratory showed that samples taken from the upper elevations in B-172 

and B-174 (from elevations 96 to 90 ft) contained non-clay material. A second plot, with a 

maximum erosion rate of 50 mm/h, is provided in figure 22 that more clearly shows the range of 

data points from the less erodible samples. Figure 23 shows the erosion data in log-log scale,

along with the erosion categories proposed by Briaud et al. (2011).
21

Source: FHWA.

1 mm/h = 0.039 inch/h.

Figure 21. Graph. Erosion rates from ESTD testing.

Source: FHWA.

Figure 22. Graph. Erosion rates from ESTD testing under 50 mm/h.
22

Source: FHWA.

Note: Dashed lines represent boundaries of erosion categories.

Figure 23. Graph. Erosion rates of samples with corresponding erosion categories.

The Geotechnical Laboratory classified samples ST-11 from B-172 and ST-23 from B-174 as 

poorly graded sand with silt (SP-SM) per the USCS. Both soils were nonplastic. The boring logs 

suggested that the samples were collected from the Black Creek Formation, where ST-11 was 

designated as a clayey soil type and ST-23 was designated as a silty or clayey gravel and sand 

type. On the other hand, the CPT soundings (figure 18) suggested the soils at both locations were 

sand mixtures with layers of silt mixtures. Photos of samples ST-11 and ST-23 are shown in 

figure 24.
23

All photos source: FHWA.

A. ST-11, B-172. B. ST-23, B-174.

Figure 24. Photos. ST-11 and ST-23 specimens.

Although samples ST-11 and ST-23 were classified as sand with silt, it is unclear if samples 

ST-9 and ST-10 from B-172 and ST-22 and ST-24 from B-174 were similar materials or if there 

were clay layers present in those samples. Sample ST-9 from B-172 did exhibit some cohesive 

behavior during erosion and shear testing with low erosion rates that corresponded with clay. 

Although FHWA researchers believed that some of these samples may have contained some 

clay, specifically ST-9 and ST-10, they did not have confidence in any of the results from the 

unclassified samples and recommended against including that data in the critical shear stress

analysis.

Sample ST-2 from B-157, however, came from a potential clay layer. No soil classifications 

were conducted directly on that sample, but the Geotechnical Laboratory did classify sample 

ST-3, which was directly beneath it. Sample ST-3 was classified as a highly plastic clay (CH) per 

the USCS. The B-157 boring log suggested that the layer, which contained both samples, was 

gray, highly plastic, fine sandy, and silty clay and was designated as a clayey soil type. On the 

other hand, the CPT soundings suggested that the layer was silt mixtures with layers of clay and 

sand mixtures (figure 18). Figure 25 shows a photo of sample ST-3.

Source: FHWA.

Figure 25. Photo. ST-3 B-157 specimen.
24

Based on the results of the soil classification of ST-3 and the CPT data from B-157, FHWA 

researchers were confident that the ST-2 sample tested in the ESTD was representative of the silt 

and clay layer found at elevations 80–85 ft. The erosion data from sample ST-2 testing in the 

ESTD is shown in figure 26. Figure 27 shows the ST-2 erosion data in log-log scale, along with 

the erosion categories proposed by Briaud et al. (2011). Because of the data scatter in figure 26

and figure 27, it was very difficult to fit the data with a single curve to calculate a critical shear 

stress value. Therefore, alternative methods were considered to derive a shear stress distribution 

for the entire clay layer instead of only finding a deterministic value of the critical shear stress.

Source: FHWA.

Figure 26. Graph. Erosion rates from ESTD testing of ST-2.
25

Source: FHWA.

Note: Dashed lines represent boundaries of erosion categories.

Figure 27. Graph. Erosion rates from ST-2 with corresponding erosion categories.

CRITICAL SHEAR STRESS DISTRIBUTION

Because the ESTD can produce a wide range of flow rates and, thus erosion rates, a probabilistic 

analysis method was considered to determine the critical shear stress value along with a 

confidence interval, typically expressed as a coefficient of variation (COV). FHWA considered 

multiple methods to determine these values. Before any method was applied, the data was broken 

down into smaller time windows. During the 10-min erosion test, it is not unusual for the soil to 

erode at different rates throughout the run. The simplest method to interpret the erosion rate was 

to calculate a best fit line across the entire length. However, if the data are broken down into 

smaller time windows—varying from 1 to 4 min, with overlaps of up to 50 percent—more data 

points are generated that better represent the full range of erosion seen in the original 10-min run.

Two methods were considered for determining the critical shear stress of the soil using a 

probabilistic analysis. The first method was the bin method. The set of data points was separated

into a series of horizontal bins containing a certain range of erosion rates, where the total number 

of bins was equal to the log-base 2 value of the number of data points. Each bin contained an 

equal number of data points. For each bin, the mean, standard deviation, and COV were 

calculated for both erosion rates and shear stress values. Then the mean values were plotted and 

fitted using the erosion power function to find a resulting critical shear stress.
26

The second approach was the bootstrapping method (Stine 1989). In this method, the data are 

plotted on a log-log plot, and from the total number of data points, a subset is selected at random. 

From this subset, a log-log best fit line is calculated. This process is repeated N number of times, 

anywhere from 5,000 to 50,000. For this method, a nonzero erosion rate must be selected to find 

the critical shear stress value, typically 0.1 mm/h (0.0039 inch/h), which matched the definition 

used by Briaud et al. (2011). The intersect of the best fit lines at the selected erosion rate results 

in a distribution of critical shear stress values, from which a COV can also be calculated. Both 

methods were tested by FHWA to evaluate the optimal way to determine the probabilistic 

distribution of critical shear stress values for a tested soil. The bootstrapping method was 

eventually selected because it generated more reliable and reproducible distributions from the 

erosion datasets.

The first step of the bootstrapping method was to apply a power function to describe the erosion 

function of the cohesive soil tested. The erosion rate is given by equation 2.

(2)

Where:

ė = erosion rate.

τ = shear stress.

ka, kb = equation constants.

By taking the logarithm of both sides, equation 2 converts to a linear relationship (as shown in 

equation 3), and a linear best fit was applied to the data to obtain constants ka and kb. After fitting 

both constants, the critical shear stress was defined as the shear stress when the erosion rate 

equaled 0.1 mm/h.

(3)

The erosion data was divided into 4-min time windows, with 2-min overlaps between adjacent 

windows. A linear best fit function was applied to each window, where the slope of the linear fit 

was equal to the erosion rate. For very low flows, where erosion may not be detectable by the 

laser, a lower boundary for erosion was set at 0.016 mm/h (6.3×10-4inch/h), which was based on 

the resolution of the laser scanner. In total, 314 erosion data points were collected (figure 28). 

The solid line represented the linear fit function for the entire dataset. With the corresponding 

fitted constants ka of 0.0091 and kb of 0.6611, the critical shear stress was calculated to be 

37.4 Pa (0.785 psf).
27

Source: FHWA.

Figure 28. Graph. Logarithmic best fit of the erosion data.

The second step of the bootstrapping method was to randomly select a subset of data points from 

the total set and apply the logarithmic best fit equation. A script randomly selected 100 out of the 

314 total data points and then applied the best fit function to calculate a corresponding critical 

shear stress. This process was then repeated 50,000 times to get a distribution of critical shear 

stresses (figure 29). The two dashed lines in figure 29 represent the 95-percent confidence limits 

on the linear fit of the mean critical shear stress. The solid line is the mean linear fit of the 50,000 

iterations. Figure 30 plots a histogram of all 50,000 critical shear stresses. The mean value of the 

critical shear stresses was 38.7 Pa (0.813 psf), which was within a few percent of the value 

calculated previously for the entire 314-point dataset. The standard deviation was 12.9 Pa 

(0.271 psf), and the COV was 0.33. Both mean and COV were used as the critical shear stress 

distribution parameters. Since the erosion rates follow a lognormal distribution, the distribution 

of the critical shear stress was also assumed to be lognormal.
28

Source: FHWA.

Note: Dashed lines represent the 95-percent confidence interval on the mean linear fit.

Figure 29. Graph. Bootstrapping technique showing 50,000 linear fittings.

Source: FHWA.

Figure 30. Graph. Histogram of the 50,000 critical shear stresses.

SUPPLEMENTAL EFA TESTING

Shelby tube samples were shared with the Geotechnical Laboratory to conduct additional erosion 

testing with the EFA to supplement the ESTD testing and help validate the shear stress results. 

Laboratory technicians first tested ST-3 from B-157, ST-11 from B-172, and ST-23 from B-174.

After determining that samples ST-11 and ST-23 contained poorly graded sand with silt, FHWA 

researchers selected samples ST-15 from B-172 and ST-29 from B-174 for additional testing in 

the EFA. These two samples came from silt/clay layers similar to ST-3, according to the CPT 

data shown in figure 18, and could help establish confidence that the clay layer existed across the 

site.
29

Sample ST-15 from B-172 was classified as a highly plastic clay (CH) per the USCS. The B-172 

boring log suggested the sample was collected from the Black Creek Formation and was 

designated as a clayey soil type. On the other hand, the CPT sounding suggested the soil at the 

specimen location was predominantly silt mixtures and clays. Sample ST-29 from B-174 was 

classified as a low plasticity clay (CL) per the USCS. The B-174 boring log suggested that the 

sample was also collected from the Black Creek Formation and was designated as a clayey soil 

type. On the other hand, the CPT soundings suggested that the soil at the specimen location was 

predominantly silt mixtures and clays. Figure 31 shows photos of the surfaces of samples ST-15 

and ST-29. Table 6 summarizes the geotechnical results from the five samples classified by the 

Geotechnical Laboratory.

Source: FHWA.

A. ST-15, B-172. B. ST-29, B-174.

Figure 31. Photos. ST-15 and ST-29 specimens.

Table 6. Summary of geotechnical results for NCDOT soil specimens.

Sample ID

Elevation

(ft)

WC

(percent)

Gravel*

(percent)

Sand**

(percent)

Fines***

(percent)

D50

(inches)

ST-3 B-157 83.8–81.8 24.5 0.0 9.2 90.8 1.5E-05

ST-11 B-172 92.0–90.0 25.5 0.0 93.5 6.5 0.0138

ST-15 B-172 84.0–82.0 24.4 0.0 14.2 85.8 3.7E-05

ST-23 B-174 96.0–94.0 19.1 0.4 91.2 8.8 0.0164

ST-29 B-174 84.0–82.0 23.3 0.0 14.0 86.0 5.5E-05

*Particle sizes > 0.1874 inches.

**Particle sizes between 0.1874 and 0.0029 inches.

***Particle sizes < 0.0029 inches.

WC = water content.
30

Table 6 continued. Summary of geotechnical results for NCDOT soil specimens.

Sample ID

LL

(percent)

PI

(percent)

Classification qu

USCS AASHTO* (tsf)

ST-3 B-157 55 35 CH A-7-6 1.9

ST-11 B-172 NP NP SP-SM A-2-4 >4.5

ST-15 B-172 53 32 CH A-7-6 2.85

ST-23 B-174 NP NP SP-SM A-1-b 4.11

ST-29 B-174 47 28 CL A-7-6 2.75

LL = liquid limit; PI = plasticity index; AASHTO = American Association of State Highway and Transportation 

Officials; NP = nonplastic; tsf = tons per square ft.

*Per AASHTO M 145 (1995).

The five samples were tested in the EFA, which is shown in figure 32. The soil in the Shelby 

tube was first extruded, trimmed, and then placed and tested in the EFA under a range of water 

velocities. The test was run until either 0.4-inches of the specimen had eroded or 60 min had 

elapsed, whichever occurred first. Then, the sample was removed from the device, and the 

surface roughness of the specimen was measured using a laser texture scanner. The hydraulic 

shear stress at the water-soil interface was calculated for each speed using the Moody chart

(Moody 1944). The relationship between erosion rate and hydraulic shear stress/velocity for each 

specimen (i.e., the erosion function of the soil) was then determined (Figure 33 and figure 34, 

respectively). Table 7 shows each specimen’s critical shear stress and velocity (i.e., the 

stress/velocity at which erosion initiates). Note that the ST-29 B-174 specimen did not erode 

under the maximum possible velocity of 19.0 ft/s in the EFA device, so its critical shear stress 

and velocity could not be determined.

(a) (b)

Source: FHWA. Source: FHWA.

Note: Dimensions in mm. 1 mm = 0.039 inch.

Figure 32. Sketch. (a) EFA cross section based on Briaud (1999), and (b) laboratory device

at TFHRC (Nicks et al, 2023).
31

Source: FHWA.

1 N/m2 = 0.021 psf.

EC = erosion categories.

Figure 33. Graph. EFA erosion function results, shear stress.

Source: FHWA.

1 m/s = 3.28 ft/s.

Figure 34. Graph. EFA erosion function results, velocity.
32

Table 7. EFA critical shear stress results.

Sample ID

Critical Shear Stress Critical Velocity

N/m2lb/ft2 m/s ft/s

ST-3 B-157 36.23 0.767 3.0 9.8

ST-11 B-172 4.59 0.096 1.0 3.3

ST-15 B-172 100.56 2.10 4.8 15.7

ST-23 B-174 4.24 0.092 1.0 3.3

ST-29 B-174* >132.97 >2.777 >5.8 >19.0

*Specimen did not erode; values provided were from the maximum applied EFA velocity.

These results show that the clay samples, including ST-3 from B-157 and ST-15 from B-172 fit 

in the low erodibility erosion category. Clay sample ST-29 from B-174 was not erodible by the 

EFA, so at a minimum it fits in the very low erodibility erosion category. Samples ST-11 from 

B-172 and ST-23 from B-174, which were classified as poorly graded sand with silt, had 

significantly less erosion resistance than the clay soils and fit in the moderate erodibility erosion 

category.

The critical shear stress calculated for ST-3 in the EFA was 36.23 Pa (0.761 psf), which was very 

similar to the results of the bootstrapping analysis conducted for ST-2, which had a mean value 

of 38.7 Pa (0.813 psf). The close results between the EFA and ESTD devices gave FHWA 

researchers confidence in a mean critical shear stress value of 38.7 Pa (0.813 psf) for the clay 

layer starting at elevation 83 ft. Figure 35 compares the erosion data results from the two 

devices. The samples tested in the layers between elevations 96 to 90 ft were not considered in 

the final analysis because of the uncertainties with soil classification.
33

Source: FHWA.

Note: Dashed lines represent boundaries of erosion categories.

Figure 35. Graph. Comparison of EFA and ESTD results.

NCDOT ADDITIONAL SOIL INVESTIGATIONS

NCDOT conducted additional borings in May and June 2022 to continue their geotechnical 

investigation of the site and collected dozens of SS samples to perform soil classifications. 

Boring B1-A was conducted close to B-157, and boring B3-B was near B-174. Boring B1-A 

suggested a layer of dark gray and brown, silty clay from elevation 89.5 ft to 79.5 ft. Boring 

B3-B suggested a layer of gray, silty clay from elevation 100.3 to 95.6 ft, followed by sand and 

silt layers, then a layer of dark gray and red silty clay from elevation 88.3 to 80.3 ft.

Laboratory testing conducted by NCDOT showed that SS samples recovered from these two 

borings between elevation 89 to 82 ft had similar fines content (84–89 percent) and similar liquid 

limit (LL) (53–69 percent) and plasticity index (PI) (30–42 percent) as the EFA samples 

summarized in table 6. However, for boring B3-B, the two SS samples recovered from elevation 

100 to 97 ft showed fines contents of 68.8 and 78.2 percent, LL content of 43 percent, and PI 

content of 16 and 14 percent, respectively, indicating low plasticity silt per the USCS. Cohesion 

from this silty material at this depth may explain why ST-22 from B-174 (elevation 97.5 ft), 

which visually had sand layers intermixed, produced some ESTD data points in the low 

erodibility category in figure 23. FHWA has not conducted enough testing on silts, however, to 

have confidence in those results.
34

FHWA also reviewed NCDOT laboratory testing results of SS samples collected in 2021 from 

B-157, B-172, and B-174 when Shelby tube samples were recovered at the site. These SS 

samples were collected both above and below the Shelby tube depths (detailed in figure 18). 

Percent fines values from these SS samples ranged from 65 to 81 percent, which was lower than 

the CH/CL samples tested in table 6.

Given the intermixed layers of sands and silts at higher elevations, FHWA researchers only have

confidence in applying the critical shear stress value of 38.7 Pa (0.813 psf) found in ESTD and 

EFA testing to other clays at the bridge site within the same general elevation of 83 ft and with 

similar properties, including percent fines, liquid limit, and plasticity index.
35

CHAPTER 5. TASK 3. DETERMINISTIC SCOUR ANALYSIS

In Task 1, the researchers used CFD modeling to obtain the water load, i.e., nominal shear 

stresses on the initial riverbed near the structure. In Task 2, soil erosion testing was used to 

obtain the erosion resistance, i.e., mean and distribution of the soil’s critical shear stress. In order 

to determine the resulting scour depth, decay functions were derived to describe how the 

hydraulic shear stresses decay as the scour hole deepens. Two methods are used to obtain the 

decay functions: the HEC-18 method and the analysis of scour data from flume experiments, 

including tests conducted by Annandale and Jones and FHWA (Arneson et al. 2012; Annandale 

2006).

PIER SCOUR DECAY FUNCTION DEVELOPMENT

The HEC-18 method to develop a decay function involves combining and rearranging several 

design equations (Arneson et al. 2012). Equation 4 (HEC-18 equation 7.40) describes the pier 

scour decay function along with the scour depth in terms of the stream power:

(4)

Where:

P = the stream power at any given scour depth.

Pa = the stream power of the approach flow near the stream bed.

ys = scour depth.

B = pier width perpendicular to the flow direction.

Equation 5 (HEC-18 equation 7.39) relates the approach flow stream power to the bed shear 

stress (Arneson et al. 2012):

(5)

Where:

ρ = water density.

τ = bed shear stress of approach flow.

Assuming equation 5 is still valid at the pier, then the following relationship applies:

(6)

Then, the pier scour decay function in terms of shear stress can be derived as follows:

(7)
36

The second method for developing a decay function involves analyzing the results of scour data 

from flume experiments. When the FHWA Hydraulics Laboratory conducted a similar TPF 

study for the Michigan DOT (MDOT), the hydraulics researchers performed physical modeling 

of several scaled bridge pier configurations in their flume (FHWA 2023). The MDOT project 

featured large bascule piers with sheet-pile coffer dams and a fender system. The scaled models 

included pier only, pier with sheet pile, and pier with the fender. Flume scour tests ran for 70 to 

120 hours and were paused at regular intervals to scan the bathymetry and measure the 

incremental progress of the scour hole. These scans were then imported into CFD to calculate the 

bed shear stresses in the scour hole (FHWA 2023). These data, together with data from 

Annandale and Jones (Annandale 2006) were normalized and then plotted in figure 36.

Source: FHWA.

Figure 36. Graph. Decay function from FHWA’s flume test data and Annandale and Jones 

data (Annandale 2006).

Figure 36 indicates that data from the three MDOT scaled bridge models reasonably aligned with 

Annandale’s and Jones’ results. The best fit equation through the data, also referred to as the best

fit decay function, is given by equation 8. Shan et al. (2016) used a reliability index (RI) to 

measure the reliability and accuracy of an equation. The same RI analysis was performed here to 

give the best fit decay function a reliability index of 2.0. Equation 8 was multiplied by a safety 

factor of 1.60 to get the more conservative design decay function, which is given in equation 9.

(8)

(9)
37

DETERMINISTIC PIER SCOUR ANALYSIS

The deterministic pier scour analysis was conducted once the initial riverbed shear stress, decay 

function, and critical shear stress of the clay layer were known. The design decay function 

(equation 9) was used to check whether the clay layer would be eroded under a given flow.

First, the upstream approach shear stresses, τa, for Q100 and Q500 were computed using 

equation 10 (equation 6.7 in HEC-18) with given HEC-RAS flow parameters at the approach 

cross section of station 364713 (Arneson et al. 2012; USACE 2022).

(10)

Where:

γ = unit weight of water.

y = approach flow depth.

n = Manning’s n.

V = average approach velocity.

Ku = 1.0 for SI and 1.486 for English units.

This calculation resulted in approach shear stress values of 2.67 Pa (0.056 psf) and 2.52 Pa 

(0.053 psf) for the Q100 and Q500 flows, respectively. The calculated approach shear stress was 

then converted to bed shear stress at the bent using the ratio from the design equation from 

equation 8. The approach shear stresses calculated from equation 9 can overestimate the shear 

stress compared to the results from CFD simulations because it is averaged over the entire cross 

section and relies heavily on the estimation of Manning’s n value. The results from CFD 

simulations consider the full velocity profile of the flow and compute the shear stress in the 

boundary layer of the surface. However, at the Lumber River site, the approach shear stress 

values of 2.67 and 2.52 Pa (0.056 and 0.053 psf) were in a similar range to the values calculated 

using CFD because the velocities at this site were very low.

Compared to equation 10, a CFD simulation can more accurately capture the 3D flow patterns

near the piers and abutments. Additionally, because CFD modeling of bed shear stresses at the 

piers and abutments already considers the contraction effect, no separate contraction scour 

analysis was needed. Therefore, for this study, the initial bed shear stresses calculated near the 

bents on the riverbed in Task 1 were used as the initial starting condition for the deterministic 

analysis instead of the results from equation 10. These values, originally listed in table 4, were 

5.0 Pa (0.105 psf) for the Q100 and 5.6 Pa (0.118 psf) for the Q500 flood event. These values were 

then multiplied by the safety factor of 1.6 used to get the design decay function (equation 9) from 

the best fit decay function (equation 8). This factor increased the initial bed shear stresses at the 

pier to 8.0 and 8.96 Pa (0.168 and 0.188 psf), respectively. FHWA recognizes that CFD 

simulations will not be available for all bridge designers, so future research will focus on 

determining a table of shear stress modification factors to get appropriate initial bed shear stress

values at the piers from the approach shear stress calculated using equation 10.

With initial bed shear stress values selected, the decayed shear stress was computed at 1-ft 

incremental scour depths using equation 9. These shear stresses were compared with the clay
38

critical shear stress value in the clay layer. Figure 37 reveals that at 83-ft elevation, the decayed 

shear stress was 6.23 Pa (0.131 psf) for Q500, much less than the design clay erosion resistance of 

25.8 Pa (0.542 psf), which was equal to the mean shear stress minus one standard deviation. 

Therefore, the total pier scour would stop at 83 ft under these conditions.

Source: FHWA.

Figure 37. Graph. Deterministic pier scour analysis using decay function and clay 

resistance for Q500.

ABUTMENT SCOUR DECAY FUNCTION DEVELOPMENT

Decay functions for abutment scour were once again developed using design equations from 

HEC-18 (Arneson et al. 2012). Equation 11 (HEC-18 equation 8.7) gives an alternative 

clear-water contraction scour equation:

(11)

Where: 

yc = flow depth, including contraction scour.

γ = unit water weight.

τc = critical shear stress of the riverbed material.

n = Manning n.

q2 = unit discharge in the constricted opening accounting for nonuniform flow distribution.

Ku = 1.0 for SI and 1.486 for English units.

Together with HEC-18 equation 8.3, the abutment scour can be calculated as follows (Arneson 

et al. 2012):
39

(12)

Where: 

ymax = flow depth, including abutment scour (ymax = yabut + y0).

y0 = flow depth prior to abutment scour.

α = scour amplification factor.

If equation 11 is rearranged and combined with equation 12, the abutment scour decay function 

in terms of shear stress can be obtained:

(13)

When the flow parameters at the bridge cross section are extracted from the HEC-RAS model, 

equation 13 can be used to calculate the initial riverbed shear stress at the abutment (USACE 

2022). However, similar to the pier scour decay function, the nominal abutment shear stress 

value from the CFD analysis was used as the initial condition for the decay function instead of 

equation 13. According to table 4, the highest nominal shear stress value for either abutment 

occurs at the left abutment during Q100, with a value of 17.8 Pa (0.374 psf). Therefore, the worse 

abutment scour scenario happens at Q100 at the left abutment and was used for the deterministic 

abutment scour analysis.

DETERMINISTIC ABUTMENT SCOUR ANALYSIS

With the initial bed shear stress value selected at the abutment, the decayed shear stress was 

computed at 1-ft incremental scour depths using equation 13. The unit discharges and flow depth 

were taken from the HEC-RAS model from a cross section upstream of the bridge and another at 

the bridge opening (USACE 2022). The scour amplification factor, α, was determined by 

plugging the ratio of q2/q1 into figure 8.10 in HEC-18, which resulted in an amplification factor 

of 1.75 (Arneson et al. 2012). Figure 8.10 in HEC-18 represents the scour amplification factor 

for wing-wall abutments and live-bed conditions; however, it was selected because it represented

the worst-case scenario. When all the values were entered into equation 13, the result was 

19.3 Pa (0.405 psf), which was slightly reduced to match the CFD initial condition results of 

17.8 Pa (0.374 psf) from table 4.

The decaying shear stress load was compared against the clay critical shear stress value in the 

clay layer. Figure 38 reveals that at 83-ft elevation, the decayed shear stress was 4.3 Pa 

(0.090 psf) for Q100, much less than the design clay erosion resistance of 25.8 Pa (0.542 psf). 

Therefore, the abutment scour would stop at 83 ft under these conditions. This analysis assumes 

that the same clay layer exists at each abutment. Additional borings or CPTs may be needed to 

confirm the extent and variability of the clay layer at the abutments.
40

Source: FHWA.

Figure 38. Graph. Deterministic abutment scour analysis for Q100.

SCOUR ANALYSES USING HEC-18 CLAY SCOUR EQUATIONS

HEC-18 contains an alternative methodology to decay functions to calculate scour in cohesive 

soils (Arneson et al. 2012). Equation 14 (HEC-18 equation 6.6) calculates the ultimate 

contraction scour depth in cohesive soils if the critical shear stress of the cohesive soil is known.

(14)

Where:

ys-ult = ultimate contraction scour depth in the cohesive soils.

y1 = upstream average flow depth.

V2 = average flow velocity in the contracted section.

ρw = water density.

g = gravitational constant.

Similarly, equation 15 (HEC-18 equation 7.35) calculates pier scour in cohesive soils if the 

critical velocity of the cohesive soil is known (Arneson et al. 2012).

(15)

Where:

ys = ultimate pier scour in the cohesive soils.

K1 = correction factor for pier nose shape.
41

Where:

ys = ultimate pier scour in the cohesive soils.

K1 = correction factor for pier nose shape.

K2 = correction factor for the angle of attack of the flow.

a = projected pier width.

V1 = flow average velocity directly upstream of the pier.

Vc = critical velocity for initiation of erosion of the cohesive soils.

Since scour in cohesive soils happens much slower than in noncohesive soils, HEC-18 also gives 

time-rate scour equations to adjust the ultimate scour depths (Arneson et al. 2012). However, 

assigning flow durations to specific flood events is currently challenging for designers. 

Therefore, table 8 only summarizes the ultimate scour depths to compare results from the 

deterministic scour analyses and does not consider the time-rate scour. Also, the calculations in 

table 8 assumed that the entire soil profile consists of the same clay material. Table 8 reveals that 

HEC-18 clay scour equations result in slightly deeper ultimate scour depths than deterministic 

scour analyses using the decay functions.

Table 8. Flow variables and scour calculations of ultimate contraction scour and pier scour 

in cohesive soils.

Flow

y1

(ft)

V2

(ft/s)

τc

(Pa)

V1 

(ft/s)

Vc

(ft/s)

a

(ft)

ys

*

(ft)

ys-ult

**

(ft)

ys-total

***

(ft)

Total Pier 

Scour 

Elevation

(ft)

Q100 20.15 4.45 38.6 4.45 6.2 3 17.1 1.4 18.5 81.5

Q500 22.12 5.01 38.6 5.01 6.3 3 20.1 2.3 22.4 77.6

*Ultimate pier scour depth.

**Ultimate contraction scour depth.

***Total combined pier and contraction scour depth.

43

CHAPTER 6. TASK 4. LONGITUDINAL SCOUR ANALYSIS AT VERTICAL WALL 

AT DOWNSTREAM BEND

Downstream of the I–6064/I–95 bridge, a meander bend of the Lumber River is directly adjacent 

to the proposed highway, resulting in a potential scour hazard of the embankment. NCDOT 

prefers not to excavate the channel at the bend to install buried riprap countermeasures and 

instead plans to install a vertical sheet pile wall with lightweight fill and deadman tiebacks next 

to the highway. The CFD model developed for Task 1 also calculated bed shear stresses for the 

bend area, which were used to evaluate potential lateral and vertical scour. Figure 39 shows the 

shear stress distribution at the bend for the Q500 flood event. Table 9 lists the surface average and 

maximum shear stress values for four different design floods. The maximum value of 12.5 Pa 

(0.263 psf) was much smaller than the design critical shear stress of 25.8 Pa (0.542 psf) for the 

clay soil layer. However, FHWA did not test any samples from the bend in the ESTD or the 

EFA. If NCDOT is confident that the same clay layer exists at the bend, as illustrated in figure 

40, it could potentially protect the base of the vertical wall from additional scour.

Source: FHWA.

Figure 39. Image. Shear stress distribution in the Lumber River bend.

Table 9. Shear stress values in the Lumber River bend for various flood events.

Flow 

Frequency

Surface Average 

Shear Stress (Pa)

Maximum Shear 

Stress (Pa)

Q25 2.3 8.1

Q50 2.8 7.4

Q100 2.8 8.8

Q500 2.6 12.5
44

Source: FHWA.

τB.Max = maximum shear stress in the bend.

Figure 40. Sketch. Profile sketch for the Lumber River bend.

In March 2021, NCDOT returned to the bridge site to collect an additional boring (B-28) at the 

bend location. The B-28 boring log suggested that at elevation 101.3 ft was a layer of gray, sandy 

clay with shell fragments and was part of the Black Creek Formation. On the other hand, an 

analysis of a split-spoon sample collected at 91.3 ft, conducted by an NCDOT contractor,

classified it as a clayey sand (SC) per the USCS. The B-28 boring log then suggested that at 

elevation 82.8 ft was a layer of light gray, clayey sand. No samples were collected from this 

layer for indexing. At elevation 72.8 ft, the B-28 boring suggested that there was a layer of gray 

and tan silty sand. Again, no samples were collected from this layer for indexing. Three Shelby 

tube samples were collected from the clayey sand layer from elevation 101.3 to 82.8 ft but were 

not shipped to or tested by FHWA. Given the general descriptions from the boring log, NCDOT 

is cautioned against strictly applying the critical shear stresses found for the CH and CL layers at 

the bridge location for this bend location in determining appropriate scour depths.
45

CHAPTER 7. SUMMARY

The proposed NCDOT I–6064/I–95 bridge replacement project over the Lumber River provided 

an excellent candidate for the TPF study to advance NextScour research. FHWA researchers in 

the Hydraulics and Geotechnical Laboratories conducted SRH-2D and 3D CFD modeling, soil 

erosion testing, and soil indexing.

Hydraulic modeling using 2D and 3D models found an initial representative bed shear stress of 

5.0 Pa (0.105 psf) for Q100 and 5.6 (0.118 psf) for Q500 at bent R4, which considered the skew 

angle as the flow came around the left abutment. These values were used as the worst case shear 

stress when calculating pier scour. The initial bed shear stress values for the abutments were 

17.8 Pa (0.374 psf) and 16.5 Pa (0.347 psf) for the Q100 and Q500 discharges, respectively.

Soil erosion testing was conducted to determine the critical shear stress of the clay layer. Based 

on the boring logs, it was initially assumed that the clay layer started relatively close to the 

surface at an elevation of 96 ft, but geotechnical indexing revealed that soil layer contained 

mostly sand and silt mixtures. However, the layer at elevation 83 ft was classified as clay and 

was used for erosion resistance analysis. ESTD and EFA testing found the mean critical shear 

stress of the clay was 38.7 Pa (0.813 psf). A deterministic scour analysis was accomplished by 

comparing the decayed shear stress results using the decay functions to the erosion resistance of 

the clay layer. The decay functions predicted that both pier and abutment scour would stop at the 

clay layer at elevation 83 ft. Additionally, contraction and pier scour equations in cohesive soils 

from HEC-18 were applied to calculate the ultimate contraction and pier scour depths and the

resulting ultimate total pier scour depth, although without accounting for the time rate of scour

(Arneson et al. 2012). The results from these calculations produced slightly deeper scour depths

than the results from the decay functions.

The CFD simulation analysis also included a meander bend downstream of the bridge that was 

adjacent to the highway and could potentially cause longitudinal scour along the roadway. The 

maximum bed shear stress value found by modeling at this location was 12.5 Pa (0.263 psf), 

which was smaller than the design shear stress of the clay layer. However, FHWA did not test 

any soil at this location, and whether the clay layer extends to this area of the project site is 

unclear.

This study result provides NCDOT with a research tool when conducting bridge foundation 

design. It also demonstrated NextScour could significantly improve the accuracy of bridge scour 

estimates. However, future monitoring of the bridge site is recommended to verify the scour 

predictions after a flood event.

47

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Recyclable

Recommended citation: Federal Highway Administration, 

NextScour Case Study: The I–6064/I–95 Bridge 

Replacements Over the Lumber River in Lumberton, NC 

(Washington, DC: 2024) https://doi.org/10.21949/1521472