Seepage erosion in layered stream bank material

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1 See discussions, stats, and author profiles for this publication at: Seepage erosion in layered stream bank material Article in Earth Surface Processes and Landforms September 2009 Impact Factor: CITATIONS 23 READS 59 3 authors, including: Nick Lindow North Carolina State University 4 PUBLICATIONS 36 CITATIONS Robert Evans North Carolina State University 39 PUBLICATIONS 497 CITATIONS SEE PROFILE SEE PROFILE All in-text references underlined in blue are linked to publications on ResearchGate, letting you access and read them immediately. Available from: Nick Lindow Retrieved on: 09 May 2016

2 EARTH SURFACE PROCESSES AND LANDFORMS Earth Surf. Process. Landforms 34, (2009) Copyright 2009 John Wiley & Sons, Ltd. Published online in Wiley InterScience ( Seepage erosion in layered stream bank material Nick Lindow, 1 * Garey A. Fox 2 and Robert O. Evans 1 1 North Carolina State University, Biological and Agricultural Engineering, Raleigh, North Carolina, USA 2 Department of Biosystems and Agricultural Engineering, Oklahoma State University, Stillwater, Oklahoma, USA Received 30 January 2008; Revised 26 May 2009; Accepted 23 June 2009 *Correspondence to: Nick Lindow, Biohabitats Inc., Baltimore, MD, USA. nlindow@biohabitats.com ABSTRACT: Current stream restoration practices often require anthropogenic manipulation of natural field soils to reconstruct stream banks in the absence of stabilizing vegetation. For this study, researchers conducted laboratory experiments on reconstructed, non-vegetated stream banks with layered soils experiencing seepage. The objective of the study was to determine the effect of seepage, pore water pressure, and bank geometry on erosion and bank stability of layered streambanks. The experimental design consisted of an intermediate-size soil lysimeter packed with a sandy clay loam top soil and an underlying fine sand layer at three bank slopes (90 o, 45 o and 26 o ). Shallow groundwater flow and seepage resulted in bank failure of geometrically stable banks. Pop out failures, liquid deformation, and piping were all observed failure mechanisms in the underlying sand material, dependent on the bank angle. Groundwater seepage processes created small-scale failures of the underlying sand leading to larger-scale failures of the overlying sandy clay loam. The underlying sand layer eroded according to the initial bank angle and change in overburden loading. The overlying loam layer failed along linear failure planes. The gradually sloped bank (i.e. 26 o slope) failed faster, hypothesized to be due to less confining pressure and greater vertical seepage forces. Researchers analyzed the laboratory experiments using the Bank Stability and Toe Erosion Model, version 4 1. The model calculated an accurate shear surface angle similar to the failure angle observed in the lysimeter tests. The model predicted failure only for the undercut 90 bank slope, and indicated stable conditions for the other geometries. Steeper initial bank slopes and undercut banks decreased the bank factor of safety. The observed failure mechanisms and measured saturation data indicated an interaction between overburden pressure, seepage forces, and bank slope on bank stability. Future bank stability modeling would benefit by incorporating lateral seepage erosion and soil liquefaction prediction calculations. Copyright 2009 John Wiley & Sons, Ltd. KEYWORDS: shallow groundwater flow; sapping; piping; stream restoration; bank stability Introduction Stream restoration often requires in situ reconstruction of stream banks in layered soils. Designers are faced with the difficult decision of balancing observed geomorphology, expected fluvial forces, and complex soil mechanics under highly variable moisture conditions in the riparian area to recommend a stable stream bank slope for restoration. During construction, the stabilizing forces of vegetation are usually absent. Knowledge of the driving and resistive forces imposed on stream banks is vital for understanding bank stability, sediment transport, stream evolution, and channel widening processes. Ullrich et al. (1986) found that hydraulic conductivity, capillary forces, slope of underlying sand seams, and pore water pressure in tension cracks to be the most significant forces in bank stability. The flood hydrograph, while significant, was less important in explaining bank stability. Other researchers found bank stability to be highly responsive to the discharge hydrograph of streams and therefore fluvial and confining forces. In a bankfull event, the stream banks are saturated during the rising limb of the hydrograph. During the falling limb, the banks lose the confining hydrostatic pressure of the water, and rotational failure can occur. The combination of a low stream stage and high water table lowers bank stability. Detailed analyses of bank stability in semi-arid regions of the western USA have shown that long-term moderate to high flows tended to saturate stream banks and removed matrix suction (Simon et al., 2002). In the Sieve River in Italy, Casagli et al. (1999) reported that a small reduction of matrix suction resulted in a dramatic reduction in the bank factor of safety. The falling limb of the hydrograph was the most susceptible period for failure. Numerical models have been constructed to assist with analyzing the impacts of some of these forces. For example, the USDA-ARS Bank Stability and Toe Erosion Model (BSTEM) Version 4 1 (Simon et al., 2000; Cancienne et al., 2008) is a steady-state bank stability model that calculates a bank factor of safety for multi-layer stream banks. The model focuses the stability analysis on a two-dimensional wedge of soil within a vertical cross-section of the stream bank with steady-state forces broken up into driving and resistive forces. While focusing the analysis on undercutting due to fluvial scour, the current version of the model lacks mechanisms for groundwater seepage, defined as shallow groundwater that flows from stream banks. Piping is sediment erosion due to

3 1694 N. LINDOW, G. A. FOX AND R. O. EVANS macropore flow, while sapping describes bank collapse due to seepage gradient forces or particle mobilization and undercutting (Fredlund et al., 1978; Springer et al., 1985; Howard and McLane 1988; Hagerty, 1991; Hey, 2005; Fox et al., 2006, 2007; Wilson et al., 2006). Piping and sapping are significant factors in the formation of regional drainage systems, and the occurrence is distributed throughout the world, most commonly in alluvial soil deposits with natural layering where flow concentrates in pervious strata (Hagerty, 1991). Piping results in sediment displacement by subterranean flow exiting at the bank surface. A significant hydraulic gradient at the exit location can create enough seepage force to lift and entrain sediment particles, overcoming tractive and gravity forces. Displaced sediment particles are more susceptible to erosion (Hagerty, 1991). Stream bank erosion, whether through piping, sapping or fluvial erosion, can significantly contribute to stream sediment, in some cases up to 90% total sediment yield (Rosgen, 2006). The current state of the science requires further research on bank instability and sediment transport due to seepage, and the exact mechanisms of bank failure uniquely due to seepage. The role of seepage on bank stability has been previously studied at two stream sites in Mississippi, where erosion due to seepage was found to form cavities in stream banks with layered soils of differing conductivity (Fox et al., 2006, 2007; Wilson et al., 2006). Highly permeable soils eroded out from stream banks and caused overhanging soil layers to fail. The stream banks collapsed due to undercutting and cantilever failure of the cohesive layers. The use of soil lysimeters in laboratory experiments further clarified the role of sediment transport, undercutting, and soil moisture content during failure. A need exists for additional studies at sites where seepage processes are important. At a stream restoration site in eastern North Carolina, USA, stream banks in a fluviomarine, layered soil were found to slump despite gradual side slopes during post construction monitoring. A final stable side slope was observed at a mean bank angle of 16 (Lindow, 2007). The instability was attributed to a lack of bank vegetation, a bed elevation in fine sand, and significant shallow groundwater seepage through the stream banks (Lindow, 2007). Figure 1 illustrates the localized bank slumping observed in the field, with noted tension cracks, bank failure, and lack of stream bank vegetation. The local groundwater hydraulic gradient and frequent bank inundation were Figure 1. Bank failure due to seepage undercutting and slumping. The photo was taken from a restoration site in Craven County, North Carolina, where a significant shallow groundwater gradient created localized bank instability on portions of the re-constructed channel. significantly correlated to stream bank deformation (Lindow, 2007). In the humid climate of eastern North Carolina, the regional hydrological conditions exacerbate the likelihood of seepage erosion on stream banks. The majority of excess rainfall travels as surface runoff in poorly drained and/or saturated soils (Dunne et al., 1975; Kirkby, 1975; Evans et al., 2000). However, soils directly adjacent to streams are typically better drained, and the majority of excess rainfall infiltrates and travels as shallow groundwater flow toward the stream (Heath, 1980). Fluviomarine soil deposits are typically highly weathered material, but fairly youthful ( years). Mineralogy can be mixed, and structure development low. Fluvial or fluviomarine deposited soils often exhibit distinct layers of variable transmissivity (Goodwin, 1989). Seepage flow can concentrate in the layers with high transmissivity, which can lead to piping and sapping at the stream bank. The monitoring results prompted the current study, in which we used laboratory methods to analyze a two-dimensional cross-section of the stream bank under seepage forces to determine the effect of side slope on bank stability. This paper describes seepage flows, soil moisture conditions, and failure mechanisms for three laboratory experiments on reconstructed eastern North Carolina stream banks with layered soils subject to lateral seepage. The major implication of this research is that bank slope stability design based only on fluvial hydraulics analysis may fail due to groundwater instability mechanisms. On layered banks, groundwater seepage processes may be concealed as small-scale failures of underlying less cohesive layers that can lead to larger-scale failures of overlying streambank sediment. These laboratory experiments suggested that the failure mechanism was dependent on bank angle. The results are an important component in furthering bank stability modeling and improving the stream restoration practice. Materials and Methods A laboratory scale soil lysimeter was used to mimic stream banks of a restored channel in Craven County, North Carolina. The site was a previously drained agricultural field that was restored using both natural channel and analytical design techniques to include 488 linear meters (1600 linear feet) of meandering stream and wetland habitat. The restoration site was in the Lower Coastal Plain region, where site soils consisted of layered fluviomarine deposited sediment. The soils were similar in characteristics to a Grifton soil type (United States Department of Agriculture, 2007). The epipedon was a loamy sand (0 15 cm), underlain by a massive sandy clay loam horizon from 15 to 60 cm with low hydraulic conductivity. The deep soil horizon consisted of gleyed, noncohesive, and highly conductive fine sand below 60 cm. Further information on the restoration site is available through Lindow (2007). Approximately 300 kg of sandy clay loam material and 90 kg of fine sand material was collected from a pit dug in the floodplain of the Craven County site and shipped to the Biosystems Engineering Laboratory at Oklahoma State University in Stillwater, Oklahoma, for testing. Particle size was analyzed using the hydrometer method from Gee and Or (2002). Table I includes the physical attributes of the soils used in the lysimeter tests, with the final bulk density achieved after disturbance. A direct shear test was used to measure effective cohesion and the internal friction angle of soil samples from the lysimeters. Water retention curves were measured by packing soil samples to a bulk density of 1 3 g cm 3, saturating, and draining under 0 to 600 cm H 2 O pressure conditions over porous plates. The van Genuchten (1980) parameters were

4 SEEPAGE EROSION IN LAYERED STREAM BANK MATERIAL 1695 Table I. Physical parameters for soils used in lysimeter experiments Texture Sand ( mm) % Silt ( mm) % Clay (<0 002 mm) % Bulk Density g cm 3 Sand Sandy clay loam Table II. Geotechnical parameters for soils used in lysimeter experiments Texture Hydraulic conductivity cm h 1 Residual cm 3 cm 3 Water content Saturated cm 3 cm 3 van Genuchten (1980) parameters a cm 1 n Effective cohesion, c kpa Internal friction angle, ϕ Sand Sandy Clay Loam Figure 2. The soil lysimeter used to test lateral seepage effects on bank stability. A series of 12 tensiometers were inserted through ports in the plexiglass walls to monitor pore water pressure during testing. The naming convention for the tensiometers is indicated on the figure. modeled from the results. Steady-state flow conditions were used during the lysimeter testing to calculate the hydraulic conductivity. Table II includes the soil hydraulic and geotechnical parameters of the soils used for the lysimeter testing. Fox et al. (2006) and Wilson et al. (2006) previously used the lysimeter in analysis of sediment transport due to seepage. The lysimeter was constructed of 2 cm thick clear Plexiglas material and measured 100 cm in length, 50 cm in height, and 15 cm in width. The low width dimension isolates twodimensional processes in bank stability, assuming reduced lateral soil influences and limited resistive forces at the Plexiglas/soil interface. A vertical water reservoir at one end controlled an adjustable, constant head. A porous plate with 0 32 cm diameter perforations separated the water in the reservoir from the soil. Constant head in the inflow reservoir was maintained using a Marriott bottle. The bottle was set on an HW-60 KGL scale (A&D Company, Ltd, Tokyo, Japan) in order to measure inflow in units of mass per unit time at +/ 5 g resolution. The outlet gradually constricted to allow sediment and discharge sampling. An illustration of the lysimeter and monitoring setup is shown in Figure 2. The sides of the lysimeter were instrumented with an array of 12 pencil sized tensiometers (Soil Measurement Systems, Tucson, AZ). The tensiometers were installed through 1 5 cm holes drilled in the side of the lysimeter. Measurements were located 15, 30, and 60 cm from the bank face at heights of 2 5, 9 5, 14 5, and 30 cm to allow for pore water pressure readings throughout each of the soil layers. The size of the tensiometers made it possible to make readings at depths of 15 mm from the sides of the soil column while minimizing the influence on the soil matrix. The allowable range of pore water pressure readings was 80 to 5 cm H 2 O. Pressure was recorded by individual pressure transducers (Soil Measurement Systems), connected to a CR10 datalogger (Campbell Scientific, Logan, UT). Pore water pressure was recorded at 10 s intervals. The naming convention for the tensiometers is shown in Figure 2. The soil was processed prior to testing by drying to an optimal packing moisture content of 11 14% for the loam and 12 18% for the sand. Fox et al. (2006) previously determined the optimal soil moisture for packing the lysimeter. The soil was passed through an 8 mm sieve and packed the lysimeter in 2 cm lifts. Due to the physical limitations of re-packing the soil in the lysimeter, the final bulk density was constrained to 1 3 g cm 3. The re-constructed banks were 34 cm high with 24 cm of loam over 10 cm of sand. The modeled stream banks were approximately half the size of the observed stream banks at the restoration site in Craven County. The lysimeter dimensions and desired bank slopes constrained the modeled bank height to 34 cm. The outlet from the lysimeter was controlled by a removable Plexiglas sliding plate during packing. Three lysimeter tests were performed with bank angles of 90, 45 and 26 6 to test the effect of initial side slope on bank stability and mass of erosion. The 90 slope represented an unstable slope while the 26 6 slope represented a slope similar to that observed at the field site. The slopes were achieved by cutting into the bank after packing and removing soil to the desired bank angle. During each run, measurements were collected of the inflow, outflow, pore water pressure, and sediment flux. The Marriot bottle held the pressure head equivalent to total bank height for each test. Bank failure dynamics were documented by capturing time series photographs and video. The experiments were run until failure occurred at the surface. The final geometry, failure angle, location of tension crack, and mass of eroded material were measured at the end of each run. Percentage bank saturation at failure was estimated based on the number of tensiometers

5 1696 N. LINDOW, G. A. FOX AND R. O. EVANS that measured fully saturated conditions. Each bank slope was only tested once due to time constraints, and further experimentation is recommended using similar methods. The BSTEM version 4 1 was used to further analyze failure conditions, where BSTEM is a steady-state bank stability model that calculates a bank factor of safety for multi-layer stream banks. The model accounts for up to five soil layers, positive and negative pore-water pressure, confining pressure due to stream stage, and reinforcement from vegetation. The model calculates an engineering factor of safety as the ratio of the resistive forces to the driving forces on the stream bank. The driving force is dependent on the weight component of the wedge, the depth of water in the stream, and water pressure behind the wedge (in tension cracks and positive pore water pressure). The resistant force is dependent on friction between the wedge and underlying soil and internal cohesive forces (including negative pore water pressure). Balancing these terms in the driving and resistant force equations determines a factor of bank stability, with failure mechanisms including cantilever failure of undercut banks, collapse of vertically arranged slabs, rotational slumping, and wedge failures (Springer et al., 1985; Simon et al., 2002; Hey, 2005). The three bank geometries were analyzed for an engineering factor of safety using BSTEM. Theoretically, the factor of safety at failure should be less than 1 0. BSTEM calculates the shear surface angle as the average of the mean bank angle and mean internal soil friction angle (based on direct shear tests). All other parameters were derived based on data collected from the lysimeter experiments, using the measured soil layer properties, bank geometry, and moisture conditions just prior to failure to calculate the bank factor of safety. Results Lateral seepage caused bank failure in all stream banks tested, regardless of bank angle. Failure was mainly due to pop-out failures along the toe of the slope and piping by seepage through the underlying sand layer, followed by planar slope failure of the overlying loam material. However, the mechanisms of failure varied according to initial bank angle. The intermediate sloped bank had the longest time to failure and highest percentage bank saturation. The 90 bank test and 26 6 slope test resulted in more rapid failure at lower percentage bank saturation (Table III). The following series of narratives describe each lysimeter test in detail. 90 Bank The first bank was constructed with a 90 bank angle. The sand layer was packed from 0 10 cm and the loam layer from cm. The inflow reservoir was filled to a constant head just below the top of bank at 33 cm. Initial seepage was observed from the bottom of the sand layer face after 2760 s (0 77 h). Initial sapping at the bank toe occurred after 3960 s (1 1 h), and progressively undercut to a depth of 14 cm after s (8 h), when the overhanging loam bank failed. The sand material exhibited some cohesive strength during the experiment, as the sand eroded progressively due to a series of pop-out failures, slumps, or rotational wedges as the wetting front saturated the material. The undercut reached a depth of 9 cm into the bank face, when small cantilever slices began falling out from the overlying loam layer. The loam material was stable until pore pressure in the middle loam tensiometer approached zero, and the soil lost apparent cohesive strength. A tension crack formed at 17 cm into the bank face before cantilever failure. The final average bank angle was 63. The mass of loam material that failed was 5 3 kg. Inflow was steady at 1 3 L h 1 during failure. Figure 3 illustrates the initial bank geometry and final bank geometry after failure. The bank stability modeling results are also shown with computed factor of safety for each geometry. 45 Bank The second run was constructed with a 45 bank angle and similar sand and loam layer thicknesses as the first run. The initial head in the reservoir was 32 cm, but rose to 35 cm after s (33 75 h) after equilibrium saturation had occurred without significant piping or bank slumping. The onset of erosion at the toe of the slope was observed after s (34 28 hrs). The sand material slumped from under the loam material due to sapping erosion and liquefaction. The bank slope failed in three progressively larger wedges, with complete failure to the top of bank occurring after s (35 33 h). A tension crack formed at 11 cm from the top of bank. The final average bank angle was 37. The mass of loam material that failed was 5 8 kg. The steady-state inflow recorded at failure was 1 5 L h 1. Figure 4 shows the initial and final bank geometries, with the depth of undercutting highlighted. Table III. Results of stream bank soil lysimeter tests indicating initial boundary conditions and erosion measurements Initial bank angle Time to failure 1 h Inflow at failure L h 1 Tension crack 2 cm Undercut depth 3 cm Mass of BF 4 kg Percentage saturation 5 % Time measured from initialization of flow. 2 Depth from initial top of bank. 3 Measured from toe of slope. 4 BF is the measured mass of bank failure from the loam material. 5 Calculated as the percent of tensiometers indicating positive pore water pressure at failure.

6 SEEPAGE EROSION IN LAYERED STREAM BANK MATERIAL 1697 Figure 3. Initial (top) and final (bottom) bank geometry for the 90 bank lysimeter test. The results of bank stability modeling are shown to the right of each condition. The model predicted a stable bank under initial conditions, but an unstable geometry with the undercut bank. The model does not account for seepage forces Bank The third lysimeter test had an initial bank angle of 26 6, which corresponds to a 2:1 bank slope. The soil preparation and packing procedures were similar to previous tests. The test began with a head of 33 cm in the inflow water reservoir. Seepage and erosion at the toe of the slope was noted within minutes of initialization of head. The gradient through the bank was sufficient to create a seepage zone along the upper bank and fluvial zone along the lower sand layer. Considerable sediment transport was recorded with a peak sediment flux of 142 g L 1 at the outlet flume. Steady-state inflow at failure was 23 7 L h 1 as recorded by the Marriott bottle and scale, which was an order of magnitude higher than the other test runs. This experiment consisted of less soil to restrict water flow due to the imposed slope and preferential flow around soil aggregates was also more apparent. The failure plane formed along a linear surface that extended to the top of bank without a noticeable tension crack. The final average bank angle was 21, and the mass of loam material that failed was 6 8 kg. The time at failure was 470 s (0 13 h). Figure 5 shows the initial and final bank geometries. Bank stability analysis The lysimeter results showed that the final average bank angle was similar to the recommended failure plane angle based on the calculation in BSTEM (Table IV). The failure plane and tension crack lines used in BSTEM are illustrated for each bank Table IV. Predicted bank stability factor of safety and shear surface angle from the bank stability and toe erosion model (BSTEM) Initial bank angle Final bank angle 1 Estimated shear surface angle 2 Initial geometry FS 3 Undercut geometry FS < Measured from lysimeter tests as the angle from the toe of slope to top of bank after failure. 2 The average between the mean bank angle and mean soil friction angle as calculated in BSTEM. 3 FS is the BSTEM calculated factor of safety. in Figures 3 5. A factor of safety of less than 1 was considered unstable, and less than 1 3 conditionally stable. For the 90 bank with initial bank geometry and soil moisture conditions as measured at the time of failure in the lysimeter, the indicated factor of safety was In order to account for the effect of seepage undercutting, a lateral portion of the sand layer was removed from the original bank geometry to more closely resemble the bank conditions at failure due to popouts along the toe of the slope (Figure 3). With the undercut bank geometry, measured tension crack, and soil moisture conditions based on the lysimeter results, the factor of safety decreased to less than 0 01, indicating a highly unstable bank.

7 1698 N. LINDOW, G. A. FOX AND R. O. EVANS Figure 4. Initial (top) and final (bottom) bank geometry for the 45 bank lysimeter test. The results of bank stability modeling are shown to the right of each condition. The model predicted a stable bank under both conditions, but a reduced factor of safety with the undercut bank. The model does not account for seepage forces. The model was able to predict failure of the undercut bank. A similar trend was observed for the 45 and 26 6 banks, where the initial bank conditions were predicted as stable by the model. The factor of safety dropped when the undercut bank geometry, tension crack, and moisture conditions from the lysimeter testing was entered into the model, but bank failure was not predicted. In all cases, lateral seepage through the stream banks caused failure in the lysimeter results, but the bank stability model was not able to predict bank failure. The model accurately predicted a lower factor of safety for the undercut bank conditions, but predicted failure for only the undercut, 90 bank. Discussion The failure mechanism for each bank angle was different. The 90 bank failed as a cantilever wedge due to erosion of underlying sand material arising from seepage, undercutting, and loss of apparent cohesion. The 45 bank was observed to fail as a wedge with a planar failure surface. The undercutting formed as a continuous series of pop-out failures along the toe of the slope, liquefaction of the sand material, and a large mass wasting along a linear failure plane through the upper loam material. The 26 6 bank also failed as a wedge with a linear failure plane. Significant piping and sediment erosion due to tractive forces was observed at the toe of the slope. Water infiltrating through the tension cracks accelerated bank failure. The mass of the overlying loam failure wedges decreased with initial bank angle. The depth of undercutting before mass failure also increased with initial bank angle. The trend is shown in Figure 6, and indicated an increasing overburden pressure (represented as the mass) necessary for failure with decreasing initial bank angle. The bank failure occurred faster in the more gradually sloped, 26 6 bank, which coincides with results of lysimeter experiments conducted by Wilson et al. (2007). The underlying sand layer had less confining pressure and a greater vertical component of seepage force under the gradually sloped bank condition, creating a circumstance for liquefaction, particle entrainment, and soil erosion. The vertical bank failed due to a greater driving force by the overlying soil and sapping erosion of the underlying sand layer, which is accelerated by the vertical face (Fox et al., 2006, 2007; Wilson et al., 2006). The bank undercut and the overlying loam material failed due to cantilever collapse. The intermediate bank angle may offer a balance between the vertical component of seepage force, which tended to cause particle entrainment, and the overburden pressure, which can also increase the driving force for bank instability. Liquefaction of sandy soil material in gradually sloped banks was another phenomenon observed during the lysimeter testing. The liquefaction is assumed to be due to a reduced overburden loading on the more gradually sloped banks and increased vertical seepage force. Figure 7 illustrates an example of how the pore water pressures along the bank face changed during the lysimeter testing. Figure 7B specifically shows how the pore water pressure measured in the underly-

8 SEEPAGE EROSION IN LAYERED STREAM BANK MATERIAL 1699 Figure 5. Initial (top) and final (bottom) bank geometry for the 26 6 bank lysimeter test. The results of bank stability modeling are shown to the right of each condition. The model predicted a stable bank under both conditions, but a reduced factor of safety with the undercut bank. The model does not account for seepage forces. Figure 6. Example of seepage erosion and undercutting during the 90 bank lysimeter test (left). The graph (right) shows the depth of undercutting decreased according to the initial bank angle. The mass of bank failure also decreased with initial bank angle. ing sand layer at the onset of erosion varied according to initial bank angle. Seepage erosion occurred at the toe of the slope at a lower pore water pressure for the more gradually sloped banks. This trend supported the conclusion that the increased overburden load on the sand layer decreased the likelihood of liquefaction and erosion. The gradual slope also promoted a greater vertical seepage force component. The percentage of tensiometers indicating saturation at failure was consistent between slopes (Table 3). This consistency indicated that perhaps bank saturation had the greatest effect on bank stability. The effect of saturation and positive pore water pressure has been documented as a major contributor to stream bank collapse along the River Sieve, Italy (Casagli et al., 1999; Rinaldi and Casagli, 1999) and along Goodwin Creek, MS (Simon et al., 2000). Small differences in antecedent moisture conditions between tests may have affected the results. Soil moisture in the clay soil at the beginning of testing varied from 14% for the 26 6,

9 1700 N. LINDOW, G. A. FOX AND R. O. EVANS The bank stability model predicted the shear surface angle as the average of the mean bank angle and mean soil friction angle. The final average bank angle observed after failure in the lysimeter tests closely resembled the recommended shear surface angle in the model (Table IV). The lysimeter tests indicated that the model was able to reliably predict the shear surface angle in stream banks with layered soils. Conclusions Figure 7. An example of the tensiometer readings during the 90 bank lysimeter test (location of tensiometers T3, T6, T9, and T12 shown in Figure 2). Pore water pressure increased as the wetting front advanced through the stream bank. The onset of erosion at the toe of the slope and bank failure are indicated on graph 7A. Graph 7B illustrates the change in pore water pressure observed in the underlying sand layer at the onset of erosion according to initial bank angle. Erosion at the toe of the slope occurred at a lower pore water pressure for the more gradual bank slopes, suggesting that the vertical seepage force component responsible for erosion was greater in the gradually sloped banks. Pore water pressure in the sand layer at bank failure also increased with initial bank angle. The steep vertical bank provided an overburden load that prevented liquefaction in the sand layer. 13% for the 45, to 11% for the 90 test. The sand material had initial moisture contents of 15% for the 26 6, 12% for the 45, and 18% for the 90 tests. The antecedent moisture conditions may have affected conductivity and apparent cohesion. The slight differences may explain some preferential flow around soil aggregates during the 26 6 bank test. The bank stability model, BSTEM, predicted stable bank conditions for all the initial bank geometries tested in the lysimeter. The model was also run with undercut bank geometries as observed in the lysimeter tests due to pop-out failures and sapping of the underlying sand material. The factor of safety decreased for all bank angles, with a predicted failure for the 90 bank, but stable conditions for the 45 and 26 6 banks (Table IV). The model was unable to predict the failures observed in the lysimeter tests because it does not account for lateral seepage forces. The results are further justification for integrating near bank groundwater flow processes in bank stability modeling (Darby et al., 2007; Chu-Agor et al., 2008). The lysimeter experiments indicated stream bank collapse due to seepage erosion similar to those observed at a restored stream channel in eastern North Carolina (Lindow, 2007). Seepage forces were observed to cause pop-out failures and erosion of underlying sand material in a layered stream bank, which was a critical factor in determining bank stability. The result of seepage was undercutting of the banks and failure of the overlying loam material by cantilever collapse and linear wedge slumping. The results from the lysimeter studies suggested a relationship between failure mechanism and bank slope whereby bank failure occurred faster in the more gradually sloped bank due to less confining pressure and greater vertical seepage forces. These results have important implications for future stream restoration design practices. Whenever possible, stream bank construction should be conducted during drier months, when lateral groundwater seepage is less likely to undermine stream banks. The use of slope drains during construction should also relieve some of the influence of seepage flows. Several differences were noted between field conditions and the lysimeter testing (Lindow, 2007). The measured hydraulic conductivity during the lysimeter testing was higher for the loam material than that observed in the sand material. The increased hydraulic conductivity was a result of packing procedures and low bulk density of the loam material in the lysimeters. In situ field soils had a higher bulk density (upwards of 0 3 g cm 3 ) and lower hydraulic conductivity. Limitations on the tested bank heights and slopes were also created by using the lysimeter. We assumed limited influence of the Plexiglas lysimeter walls on bank stability dynamics and thus two-dimensional analysis of the soil bank. The research would benefit from further laboratory testing and in situ measurements of seepage erosion. The stream bank model used for the study was unable to predict failure due to seepage. We recommend the inclusion of lateral seepage and near bank groundwater forces in bank stability analysis. The model did indicate a lower factor of safety for the undercut bank conditions, and predicted failure for the vertical bank with undercutting. The calculated shear surface angle reported by the model based on mean bank angle and mean soil friction angle was very similar to the observed average bank slope after failure. Acknowledgements The author would like to thank the faculty and staff in the Biosystems Engineering Department at Oklahoma State University and the Soil and Water staff from the Biological and Agricultural Engineering Department at North Carolina State University for their assistance in these laboratory experiments. The authors thank Derek Heeren and Maria Librada Chu-Agor, Oklahoma State University, Stillwater, OK, for reviewing earlier versions of this manuscript. References Cancienne R, Fox GA, Simon A Influence of seepage undercutting on the root reinforcement of streambanks. Earth Surface Processes and Landforms 33:

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