Observations of wave attenuation, scour, and subsurface...

Shore & Beach Vol. 88, No. 3 Summer 20204

Increased sea levels and growing hu-man expansion into coastal areas has resulted in greater rates of shoreline

and coastal habitat loss throughout the world. Mitigation efforts have led to increased shoreline hardening within the U.S. with over 14% of the continen-tal shoreline and 64% of all estuaries, marshes and lagoons containing some form of hardened structure as of 2015 (Gittman et al. 2016). Common hardened structures such as bulkheads, seawalls, and revetments, are constructed with materials that include metal, timber, vinyl, rock, or concrete to create a solid barrier between the shoreline and coastal waters. However, studies have shown that these “gray” protection methods result in a multitude of negative impacts including decreased biodiversity/health of intertidal ecosystems, increased passive erosion, and loss of beach front (Griggs 2005; Gittman et al. 2016).

A more modern form of coastal pro-tection known as living shorelines, has

Observations of wave attenuation, scour, and subsurface pore pressures across

three marsh restoration sill structures on a sandy bedBy

Jordan Converse,1 Meagan Wengrove,1* and Pedro Lomonaco2

1) School of Civil and Construction Engineering, 101 Kearney Hall, Oregon State University, Corvallis, OR, USA2) O.H. Hinsdale Wave Research Laboratory, 3550 S.W. Jefferson Way, Oregon State University, Corvallis, OR, USA

*Corresponding author: [email protected]

ABSTRACT With rising sea levels and more frequent exposure to extreme storms, coastlines worldwide are vulnerable to increased erosion and loss of natural marsh lands. In an effort to lessen these impacts, there is a growing practice of adapting hard or “gray” coastline protection techniques to more nature-based features that promote habitat and ecosystem health. Living shoreline marsh restorations utilize natural and nature-based materials to protect marsh shores from erosion while also allowing intertidal flushing to promote the health and diversity of the marsh. Our study investigates three types of living shoreline sill designs exposed to average and storm-energy wave conditions at varying water levels. The sills were designed to mimic constructed sills in practice (rock, oyster shell, tree root wads), but more generally vary in structure porosity and material dissipation potential. Large-scale laboratory experiments were conducted in the large wave flume at the O.H. Hinsdale Wave Research Laboratory. Wave transmission and reflection are used to demonstrate wave attenuation capability of each sill structure. Scour of the sill, bedload sediment transport rates on the seaward and shoreward sides of the sill, and sediment pore-water vertical hydraulic gradients were used to demonstrate the potential for sediment transport and liquefaction. Re-sults will contribute to understanding the effect of sill material porosity and mass on structure stability, and the effectiveness of using green living shoreline sill structures in the continued effort to establish design criteria for living shoreline implementation.

KEYWORDS: Living shoreline, root wad, oyster sill, wave attenuation, coastal protection techniques.

Manuscript submitted 1 May 2020, revised & accepted 25 July 2020.

been gaining popularity over the past decade due to its use of nature and natural materials for coastal protection and eco-system promotion. The “green” systems are multipurposed and focus on either shoreline stabilization or the creation of new ecosystems in areas where habitat was not formerly supported (Bilkovic et al. 2016). Today, best practice for living shoreline design is still being developed (Walker et al. 2011; Davis 2017; Mitchell and Bilkovic 2019). Observation and analysis presented herein are focused on marsh restoration type living shorelines. Estuarine marshes are restored with some type of vegetation, and the toe of the marsh may or may not be stabilized with a low crested sill structure as the wave climate demands. Marsh restoration focused living shorelines are typically designed based on wave climate condi-tions at the proposed construction site with low, medium, and high wave energy climates classified as areas with fetch lengths between 0.5-1 mile, 1-5 miles, and

5-15 miles, respectively (Hardaway et al. 2017). Most research has been focused on designs for a certain impact such as habitat promotion or performance in a particular geographical area (Swann 2008; Hardaway et al. 2017). For marsh restoration focused living shorelines, common designs vary from shoreline plantings of marsh grasses in low wave energy environments to plantings or pre-existing vegetation with a low crested structure placed at the marsh edge to protect existing or new marsh grasses in medium to high wave-energy climates (Walker et al. 2011). Some modern sill designs include human-made materi-als, such as bioengineered concrete shaped in the different configurations (e.g. oyster castles, reef balls), still gener-ally placed shore parallel seaward of the marsh vegetation as a low crested sill breakwater (Faherty 2011; Scyphers et al. 2011; Moody et al. 2020; Zhu et al. 2020). However, our experiment focuses on the hydrodynamics and sediment transport around the low crested sill structure at the marsh edge, and the structure’s capability for attenuating wave energy transmitted shoreward using all natural materials. The living shoreline materials chosen for the experiment are used to be representative of three different sill material densities and porosities — rock, oyster shell, and root wads. Rock sills are commonplace, as there is ample design guidelines bor-rowed from the breakwater design com-munity; however, these guidelines have led to instances of overdesigned marsh sill structures in some early projects and

Shore & Beach Vol. 88, No. 3 Summer 2020 Page 15

Figure 1. (A) Rock sill constructed of layered armored stone with a geotechnical liner between sill and bed level. (B) Oyster hybrid sill constructed of a stone core covered by oyster shell bags held in place with galvanized poultry netting to simulate established oyster colony. (C) Root wad sill constructed of three tree root wads positioned with root mat facing the seaward direction and perpendicular to the shoreline. (D) Sill cross sectional profiles at the end of testing conditions, showing that some sill settlement had occurred.

Table 1. Sill properties. Total volume and surface area were calculated based on actual sill footprint and material dimensions. Surface area Total Volume ofSill type of sill material (m2) volume (m3) Porosity solids (m3)Rock 10.07 5.38 0.48 4.59Oyster 18.73 5.18 0.77 3.48Root wad 6.35 3.85 0.8 2.61

push back from communities against the introduction of nonnative materials (Davis 2017). Oyster shell use in living shoreline sill designs has been growing in popularity in areas where oyster popula-tions persist, with experiments detailing the ability of oysters to promote habitat health and improve water quality (Hen-derson and O’Neil 2003; Whitman and Reidenbach 2012; Beseres Pollack et al. 2013; Shih and Chang 2015), as well as capacity for attenuation of wave energy using bagged oysters (Allen and Webb 2011). There are few experiments that use tree root wads as a form of coastal protection. Commonly tree root wads are used for stream bank protection and reinforcement, tree root wads are an eas-ily accessible and low-cost material in some areas of the country that provide habitat area for fish and invertebrates with minimal impact on surrounding environment and often occur naturally in northern latitude estuaries (Seal et al. 1998; Sylte and Fischenich 2000; Har-man et al. 2001). Our experiment aims to measure the physical response (wave attenuation, sill scour, sill stability) of three living shoreline sill structure design materials with varying porosities (rock, oyster sill with rock core, and tree root wads) to average and storm wave energy conditions at water level scenarios below sill crest height, at sill crest height (high tide water level), and with the sill fully submerged (high tide plus storm surge water level).

METHODS Sill design

Three sill designs, specifically, rock, oyster with a rock core, and root wads, were tested (Figure 1). The sill materials were chosen to be representative of physi-cal characteristics that are important for sill stability (e.g. weight, porosity) and potential for encouraging system habitat. The sill designs were chosen to:

1) Withstand medium-high marsh wave energy conditions,

2) Vary in sill porosity (low — rock, medium — oyster, high — root wad) and weight (medium — oyster and root wad, high — rock), and

3) Physically model hydrodynamics and sediment transport of the living shoreline low crested sill structure.

Living shoreline sills are typically de-signed to be positioned at mean lower low

water elevation with the crest elevation at mean higher high water elevation (VIMS CCRM 2017). In regions with larger tidal ranges (e.g. west and north Atlantic coasts of the U.S.), the sill is generally located at the toe of the vegetation (not located at the mean lower low water level as there is generally a fairly large expanse of mud-flats in these larger tidal range regions); the sill height is then designed so that the crest is overtopped at high water.

For our experiments, the sill crest was measured to be 0.61 m ± 0.03 m above the bed, the design height was based on the approximate average tidal range in many places along the south Atlantic and Gulf coasts; the sill height is also applicable to larger tidal range regions considering the alternative placement described. Sill heights were measured from the bed level to the crest location for the rock and oyster hybrid sills, and from the bed level to the top of the tree root trunk for the root wad sill. An additional 0.15 m of each sill was buried beneath the bed level to provide additional scour protection (i.e. total height of sill from base to crest was 0.76

m). The surface area, volume and porosity of each sill is shown in Table 1. Sill poros-ity was measured (rock, oyster hybrid) or estimated based on literature (root wad; Rafferty 2017). We estimated the effect of combined sill porosity and mass based on sill material and geometry characteristics using the sill material total surface area per sill unit volume (SA/Vsill). Material surface area is defined as the surface area of all of the material inside of the sill. The total volume of the sill was estimated based on the measured sill (L x W x H) dimensions. For the rock sill the material surface area was estimated based on the number of rocks in the sill and their average size fol-lowing Dall (1979). For the oyster sill the material surface area was estimated based on the number of oyster shells and rocks in the sill and the average surface area of an oyster shell following (Kuykendall et al. 2015). For the root wad sill, the material surface area was estimated for each root wad following (Rafferty 2017).

ROCK SILLRock living shorelines constructed of

stone can follow the same design criteria

Shore & Beach Vol. 88, No. 3 Summer 2020Page 16

as rubble mound structures as sug-gested in Miller et al. (2015). Normal environmental conditions for the sills would result in minimal overtopping as sills should only be submerged at high tide levels (Bilkovic and Mitchell 2017). Thus, the Hudson equations for non-overtopped slopped rubble mound structures were used to deter-mine the median rock size needed for the rock sill (Hudson 1974; U.S. Army Corps of Engineers 2011). A design significant wave height (Hs) of 0.56m, sill slope angle (α), density of stone (basalt) (ρs), density of water (ρw), and a stability coefficient (KD) value of 3.5 were used to determine the medium mass of stone (M50) for the sill of 12.5 kg; see Equation (1).

The stability of the stones is ensured given the reduction factors proposed for low-crested structures (e.g. Vidal et al. 1993; van der Meer and Dae-men 1994; Lomonaco et al. 2005) and the reduction of the waves due to the depth-limited conditions tested.

The storm design significant wave height was used in the calculation to determine the median stone size for the sill. Due to the small size of the sill in comparison to regular offshore rubble mound structures the calculated armor rock size was used for the entire sill construction in place of a layered ap-proach, rock mass was approximately 12.5 kg and the nominal diameter was approximately 18 cm. Rock used in the experiment was sourced from quarries in Oregon where the common quarry rock is basalt which has a density of 2,762 kg/m3 (Schlicker et al. 1978), the rock was well sorted d90/d10 = 1.25. The use of the same rock size for the entire structure in contrast of a layered rubble mound, increased significantly the permeability of the sill, reducing the development of internal pressures and reflection, both having a role in the stability of the armor.

The seaward side of sill was designed to have a 2:1 (H:V) slope to promote wave shoaling and breaking on the sill. The shoreward side of the sill had a design slope of 1.5:1 (H:V) to reduce the overall footprint of the sill (Priest III 2017). The crest width (B) of the sill was determined using Equation (2), where (W) is the primary weight and (wa) is the specific weight of the armor mate-

Equations:


Figure 2. A) Bay locations considered for wave fetch calculations. B) Mean and max fetch calculation method. Max fetch transect of nine miles was determined through analysis of longest open water distance in the predominant storm wind direction. The average distance of the max transect and subsequent transects spaced at increments of 22 and 45 degrees were used to quantify the bay mean fetch length.

rial. A layer coefficient (KΔ) of one, and a stone width (n) of three rocks were used as recommended by the Coastal Engineer-ing Manual (CEM) (U.S. Army Corps of Engineers 2011). See Equation (2).

A frontal toe was included with the sill to reduce the level of scour affecting the side of the sill exposed to incident waves. With the testing conditions located in shallow water, the construction of the toe consisted of placing two additional rows of armor layer stone that were two stones thick at the base of the sill as indicated in the CEM (U.S. Army Corps Of Engineers 2011). A geotechnical liner was placed un-der the sill as suggested by Priest III (2017) to provide structural stability and prevent upward sediment transmission through the sill as a result of propagating waves.

OYSTER SILL WITH ROCK CORE (OYSTER HYBRID)

The oyster sill was constructed using armor stone from the rock sill and oyster shell obtained from a local Oregon busi-ness (typically measuring 12 cm x 5 cm x 4 cm in size). The footprint of the sill was designed to match the crest width and slope angles of the rock sill. Oyster shell bags were constructed of 10 cm plastic mesh sleeves that were 46 cm in length per bag and secured on both ends using 16 gauge C-ring staples. Similar to the rock sill, a geotechnical liner was placed underneath the sill. The sill core was constructed using armor stone from the rock sill so that it covered one-third of the total volumetric area of the sill. Oyster

bags were then stacked on top of the stone core to match the seaward and shoreward slopes of the rock sill (Figure 1b). The sill design borrows materials (oyster shell) from breakwater designs that have proven successful in lower wave energy climates and reinforces the design using a rock core to test the effectiveness in medium wave energy conditions. A single layer of 20 gauge galvanized poultry netting was layered across the sill in the cross shore direction and secured in place with guy-wired Model 68 duckbill anchors. Anchoring the bags to the rock with this method anchored the oyster shell to the rock, as would be recommended in the interim period between living shoreline construction and growth of oyster spat into an establish oyster reef. We used plas-tic netted bags and a galvanized netting; however, biodegradable materials could be used with sill placement in a real marsh system. In practice oyster sill designs do not necessarily follow a strict engineering standard, they come in various shapes and

sizes, we chose to make the sill close to the rock sill size for comparison. Bilkovic et al. (2016) is a good reference for various oyster sill layouts in the field.

ROOT WAD SILL Root wads are commonly used in

streambank erosion control (Sylte and Fischenich 2000). Few experiments have investigated the effectiveness of root wads within coastal and marshland environ-ments. The root wads were Douglas Fir by type and were sourced from a local Oregon logging company. Due to the low density of Douglas Fir compared to water, the root wads were submerged in water for three months prior to the experiment to allow for inner pore satura-tion and to simulate the expected density in an inter-tidal environment. For the construction of the sill, three root wads were interlocked in the longshore direc-tion expanding the width of the shoreline with the root wad and tree trunk sections facing the seaward and shoreward sides,

Table 2. Experiment testing conditions. Water levels represent depth at offshore side of sill. Hs and Tp represent the significant wave height and peak period of wave spectra at the wave maker and sill locationWater Hs (m) Hs (m) just Tplevels (m) at wavemaker seaward of sill (sec) H/h kh0.33 0.25 0.19 1.69 0.76 1.45 0.56 0.41 2.89 1.69 0.90.61 0.25 0.19 1.69 0.41 1.09 0.56 0.41 2.89 0.92 0.550.91 0.25 0.19 1.69 0.27 0.92 0.56 0.41 2.89 0.62 0.41


Figure 3. Bay fetch length (top), significant wave height (middle), and peak wave period (bottom) for the 11 bays considered to inform the chosen design storm statistics. Bays are located in the continental U.S. on the East, West, and Gulf coasts. Bay locations are labeled at the top of the figure with letters over each corresponding bar, and are consistent for each panel. The dashed bars indicate the average condition for each region, and to the right, the experimental design average wave and design storm wave conditions (wave heights are shown both at the wave maker (@ WM) and just seaward of the sill (@ sill). Locations: A) Chesapeake Bay, VA; B) Delaware Bay, DE; C) Long Island Sound, NY; D) Albemarle Bay, NC; E) Galveston Bay, TX; F) Pensacola Bay, FL; G) Bay St Louis, MS; H) San Francisco Bay, CA; I) Willapa Bay, WA; J) Arcata Bay, CA; and K) Coos Bay, OR.respectively. Each root wad was secured in place with guy-wired Model 68 duck-bill anchors.

Experiment testing conditionsThe experiment was conducted in

the Large Wave Flume (LWF; 104 m x 3.7 m x 4.6 m; hydraulic piston-type wave maker) at the O.H. Hinsdale Wave Research facility located at Oregon State University. Sand sediment obtained from the Oregon coast was placed inside the flume to form the beach. The beach slope was 1:20 with a d50 sediment grain size of 0.2 mm. The sediments were placed in the LWF as part of a larger experimental

campaign, our experiment leveraged that opportunity and therefore used the exist-ing material and beach profile geometry for testing. Many marshes have sediment grain sizes equivalent to or finer than 0.2 mm; with finer grain sizes we expect sediment transport implications to be more dramatic than those presented in our analysis.

Three living shoreline sill designs and a bare beach profile with no-sill were tested over six irregular wave and water level conditions (Table 2). Two wave heights and three water levels were cho-sen to simulate a medium to high wave

energy environment during water levels below sill crest height, at sill crest height (e.g. high tide), and fully submerged wa-ter levels (e.g. high tide + storm surge). A living shoreline sill structure may be subject to any of these relative wave conditions and water levels depending on tide and surge levels during a storm. Significant wave height and peak period values were estimated at the location of the wavemaker and just before the sill, the larger waves generated broke before reaching the sill.

The low water level was 0.33 m above the bed (0.28 m below sill crest) at the offshore side of the sill and was intended to simulate a condition where waves are breaking on the seaward side of the sill, hypothesized to be the most erosive con-dition to the offshore side of the structure. The medium water condition was 0.61 m above the bed at the offshore side of the sill, slightly above the sill crest and prefer-able for intertidal flushing. The medium water level is hypothesized to subject the sill to the largest dynamic pressures as waves break over the top of the sill, potentially resulting in in liquefaction of the sediments underneath the sill leading to instability. Lastly, the high water level condition was 0.91 m above bed at the offshore side of the sill (0.3 m above sill crest). The high water level completely submerges the sill and waves break over and onshore of the sill structure, which is hypothesized to lead to both increased liquefaction and erosion on the onshore side of the sill.

Experimental wave conditions were intended to simulate average wave con-ditions and storm wave conditions. The average significant wave height and peak period were determined using a fetch length of 2 miles (3.22 km), consistent with a medium fetch wave energy en-vironment (Priest III 2017; Hardaway and Byrne 1999). Storm significant wave height and peak period were determined using wind speed data collected from the National Oceanic and Atmospheric Administration (NOAA) buoys and shore weather stations over a period of 10 years (2008-2018). Fetch driven wave statistics were calculated and averaged using 11 different marshland bays of varying fetch lengths on the East, Gulf, and West coasts of the continental U.S. (Figure 2a). Bay locations were chosen not only based upon their geography, but additionally based on availability of decadal time scale


Table 3. Sensor specifications including accuracy, range, and sampling rate. All sensor sampling rates were held constant throughout experimental testing.Device Accuracy Range Sampling rateGE Sensing (Druck) PDCR 1830 ±0.06% 0.75-600 m H2O 100 (Hz)Resistive Wave Gauge ±1.7 mm 2.37 m 100 (Hz)MTA SEATEK Transducer Array ±1 mm 0.05-4.2 m 1.0 (Hz)Nortek Vectrino (ADV) ±0.5%, ±1mm/s 2-4 m/s 100 (Hz)

or longer wind data. Individual bay wave statistics and geographical average fetch lengths and corresponding wave heights and periods are detailed in Figure 3.

The mean fetch distance was calcu-lated for each bay using a method pro-posed by Hardaway et al. (2017) where the average fetch is calculated from the mean length of five transects expanding seaward from the shoreline. The fetch distance for each transect was found by expanding a transect line from the shore-line position on one side of the bay across to another shoreline position on the op-posite side of the bay centered around the direction of the predominant storm wind direction (Figure 2b). To determine the maximum wind speeds for each bay the 10 years of wind data were individually filtered to leave only values above three standard deviations from the mean wind speed, used to identify individual storm events. Next, the design storm wind speed for each bay was defined as the mean wind speed of the three standard deviation filtered wind speeds that were calculated using a stepping function that determined the maximum wind speed over a period of 24 hours. The storm duration of all of the bays was found based on the average duration of the three standard deviation filtered wind speeds, the mean storm duration was found to be of 8 hours with a standard deviation of 2.4 hours. The storm wind event durations were checked to verify that the duration was long enough to develop fetch lim-ited waves using the fetch-limited wave equation (U.S. Army Corps of Engineers 2015) (see Equation [3]), where = time required for waves crossing a fetch length (x) influenced by a wind velocity (u) to become fetch limited.

Wind data collected from the stations varied in height and was adjusted to the standard wind speed measurement height of U10, the wind speed at 10 m above ground level. The wind speed was con-verted to the standard wind speed mea-surement at 10 m from ground level using the wind profile power law relationship (Panofsky and Dutton 1984) (see Equa-tion (4)), where Ur is the known wind speed at reference height zr, z = 10 m, α is the atmospheric stability coefficient. An atmospheric stability coefficient of 0.14, indicating neutral stability, was assumed for this experiment. Based on the analysis of the 11 bays and correction of these data, our design storm wind speed used

for the experiment was chosen to be 9 m/s with a standard deviation of 1.87 m/s.

Finally, for our calculations of sig-nificant wave height and peak period, we assumed that wind blows over a constant direction and fetch for an extended dura-tion to produce steady-state fetch limited waves following Szabados (1982). The significant wave height and peak period for each bay were calculated using the SMB (Sverdrup Munk Bretschneider) method that uses a semi-empirical for-mula to calculate the significant wave height and peak period based on wind speed and fetch length (U.S. Army Corps of Engineers 2015). The approach uses the standardized wind speed measure-ment (U10) to quantify the surface friction velocity (); see Equation (5).

Using the calculated friction velocity and bay fetch length (X), the significant wave height (Hs) and peak period (Tp) are obtained using Equations (6) and (7), respectively.

Peak periods for each wave condition were then verified against the shallow water wave growth limiting condition identified by C.L. Vincent (1985) (see Equation (8)), where d is the water depth, and g is the gravitational acceleration. The design storm significant wave height and peak period for the experiment were chosen based on these analyses, Hs = 0.56 m and Tp = 2.89 s. The experimental conditions relative to the natural design storm conditions for each of the 11 bays are presented as Figure 3.

To put the wave and water level con-ditions in context, measured significant wave height during storms at the marsh edge have been observed to be up to 0.4 m in the Chesapeake during a hurricane, up to 0.3 m during a winter cold front on the Texas Gulf coast, and up to 0.8 m during a storm in the San Francisco Bay. Additionally, surge water levels have been measured to increase water levels up to 0.5 m above high tide water level in estuarine systems. All contextual wave

heights and water levels come from per-sonal communications with researchers working in marshes around the U.S. Our experiment chose two significant wave heights that we label as “average” and ‘storm’ throughout the text (Hs = 0.19 m at the marsh edge and 0.41 m at the marsh edge, respectively); however, depending on the context of the marsh either wave height could be a storm condition for that system. The medium water level is considered to be a high tide condition for each of these systems, as low crested sills are designed to be overtopped to allow for intertidal flushing at high tide. The high water level is considered to be a high tide plus storm surge condition, which is reasonable considering field observations (i.e. high tide + 0.3 m surge in the case of the experiments presented).

Experimental setup and instrumentation

Sills and wave conditions were de-signed and tested in field scale in the Hinsdale LWF. Each trial duration was 44 minutes composed of irregular waves us-ing a TMA spectrum (Bouws et al. 1985) with significant wave height and peak pe-riod defined in Table 2 (wave maker wave heights). The TMA spectrum comes from observations in shallow water, adapted from the JONSWAP spectra, which is the spectral energy equation developed from deep water wave conditions. Sills were located 34 m from the wave maker in three water depths (Table 2).

An array of instruments were used to measure surface water elevations, veloc-ity, and sediment porewater pressures during testing (Figure 4). Six resistance wire wave gauges (RWG) were placed in the cross shore direction to capture the wave height transformation across the beach and sills. Seven Nortek Vectrino single point ADVs (Acoustic Doppler Velocimeters) were installed seaward and shoreward of the sill to capture three dimensional velocities throughout the water column. Six vertically stacked Druck 1800 pore water pressure gauges


Figure 4. Instrument set-up in the cross-shore and elevation plane. Sill footprint is outlined in gray, sand sediment in dark gray with the beach profile shown by a solid black line. Diamonds identify ADVs, RWG locations are shown by dashed lines, and pore pressure gauges by circles. Solid lines illustrate still water levels of the low, medium, and high water levels used during testing.were buried in the sediment bed seaward of the sill, under the sill front slope, and under the sill rear slope. Each stack was comprised of two pressure sensors that were vertically separated by 5.1 cm and buried 15.2 cm below the sediment-water or sediment-sill interface.

Additionally, after each run the wave maker was stopped, measurements of bathymetry were made, and the sedi-ment bed was re-set to the pre-testing conditions. A mobile multiple transducer altimeter (MTA) array, consisting of 24 horizontally mounted transducers, was used to capture the cross shore profile bathymetry. The MTA spanned 1.76 me-ters in the longshore direction and was run prior to and at the end of each trial. Accuracy, range and sampling rates of each sensor used throughout the experi-ment are shown in Table 3, instrument positions with respect to sill and water depths is shown in Figure 4.

Analysis methodsTo assess the effectiveness of each sill

type, characteristics attributed to wave at-tenuation and sill stability were examined for each wave and water level condition, including: wave transmission through the sill, incident wave reflection, bedload sediment transport at the sill front and rear slope, sill sediment vertical hydraulic gradients, and sill damage.

The wave transmission coefficient is a ratio used to describe the effectiveness of a coastal protection structure at reducing the height of incident waves reaching the shoreline. The wave transmission coeffi-cient compares the wave height transmitted across the sill, , to the incident wave height, , seaward of the coastal structure (van der Meer and Daemen 1994); see Equation (9).

Transmission coefficients are calcu-lated using RWGs mounted seaward and shoreward of the sill, where a value of one indicates no wave height reduc-tion. An RWG array located seaward of the sill location was used to capture and separate reflected waves from the offshore incident waves. Reflected wave analysis was computed using the reflec-tion analysis procedure suggested by Zelt and Skjelbreia where a one dimensional decomposition theory is used to sepa-rate measured spectra into incident and reflected wave components (Zelt and Skjelbreia 1992). Data was not collected to separate reflected waves from transmit-ted waves on the shoreward side of the sill. The wave reflection coefficient, Kr, is defined as the ratio between the reflected (Hr) and incident (Hi) wave heights (Dean and Dalrymple 1984); see Equation (10).

The velocity magnitude at each ADV location was separated into current ve-locities, U, and wave orbital velocities, uo. The current velocity magnitude is

determined by taking the time average of horizontal velocity, u, during the 44-min-ute trial. The orbital velocity magnitude is found using the root mean square (rms) velocity, and Equations (11) and (12).

A Bagnold-Bailard-Bowen type ener-getics model that calculates the bedload sediment transport rate (Q) was used to determine bedload sediment transport at the sill front and rear toes. The model uses the individual wave (uo) and current (U) components, coupled with representative friction coefficients due to waves (Kw) and currents (Kc) to calculate the sediment transport (Equation (13), Hsu et al. 2006).

Values for the friction coefficient due to waves (Kw) and currents (Kc), the fric-tion angle (ø), and the sediment transport efficiency coefficient (Єb) were obtained from Wengrove et al. (2019) because they were reformulated for small-scale transport, where coefficients in Hsu et al. (2006) were formulated for observations across a sandbar.

To quantify stability of surrounding sediments, pore water pressure data was used to quantify residual liquefaction events and scour to illustrate the structural integrity of each sill. Progressive waves cre-ate dynamic pressure loads on the seabed sediment and structure as they propagate, generating shear within the sediment. The lateral forces created by waves cause an in-crease in pore pressure within the soil that can reach levels that exceed the effective stress of the sediment, known as residual liquefaction. During residual liquefaction the pressurized pore water overcomes the interface forces between sediment grains, causing them to become unbounded (Sumer and Fredsøe 2002). To quantify the occurrence of these residual liquefaction events a threshold limit must be identi-fied. To begin, Terzaghi’s principle is used to relate the sediment effective stress (σ) and pore pressure Pw) to total stress (σ) (Terzaghi 1943); see Equation (14).

Liquefaction occurs when the effective stress equals the pore pressure, causing the total stress to be zero, so we can sepa-rate the two parts into their individual components; see Equation (15).

The effective stress is comprised of the sediment saturated specific weight (γsat) multiplied by the sediment depth (d) at the evaluation location. Likewise, the pore pressure is calculated by multiplying the specific weight of water (γw) by the


sum of d and total water head (hw); see Equation (16).

Solving (16) results in a solution that relates the change in total head to the change in sediment depth between two locations to determine the critical hy-draulic gradient (ic). The critical hydraulic gradient identifies the threshold limit for liquefaction (Holtz and Kovacs 1981); see Equation (17).

For our experiment the critical hy-draulic gradient was determined between two buried pore pressure gauges that were vertically mounted and separated by a sample length of 5.1 cm. The total head (HT) or each pressure gauge was calculated using Bernoulli’s total head expression; see Equation (18).

Pressure observations (P) collected from the pore water pressure gauges were used to quantify the pressure head (Hp) for each sensor. The elevation head was determined by multiplying the density of water (ρw) by the acceleration due to gravity (g) and the height of the sensor (h) in reference to a known datum located at the deepest pressure gauge position per stack. The velocity head (Hv) was treated as negligible due to the low seepage rates through the sediment in comparison to the magnitude of the other terms.

Sediment transport statistics, includ-ing the Shields and Sleath parameters, were examined for the potential of shear and pressure gradient induced sediment transport events. The Shields parameter compares bed level mobiliz-ing and stabilizing forces to determine the magnitude of shear driven sediment transport (Shields 1936). Our experiment uses an expression for Shields parameter introduced by Neilsen (1992) where a grain roughness factor (f2.5) is used in combination with the orbital diameter (do), specific weight of the sediment (s), and the median sediment size (d50) to determine the grain roughness Shields parameter due to wave forcing (Nielsen 1992); see Equations (19) and (20).

The Sleath parameter compares the ratio of mobilizing and stabilizing forces in an oscillatory flow to determine if plug formation, where the sediment trans-port is driven by the pressure gradient and moves as a solid block, is occurring (Sleath 1999; Foster et al. 2006). Sleath (1999) derived the plug formation (S) as the horizontal pressure gradient,

comprised of the sediment density (ρs), velocity amplitude (Uo), and angular wave frequency (ω), applied to the immersed weight of the sediment grains; see Equa-tion (21).

Finally, sill scour was quantified us-ing bathymetry profile differences for each trial determined by subtracting the pre-testing measurements from the post-testing measurements of bathymetry.

RESULTSHYDRODYNAMICS

Wave breaking location With the low water level and aver-

age wave condition trials, waves were observed to be breaking within a half-meter of the sill or at the sill edge with no overtopping and approximately 2 meters further onshore for the no-sill condi-tion. For the low water level and storm wave condition, waves were breaking in ranges that varied from on the sill front slope location to offshore locations with frequent overtopping. The medium water level and average wave conditions saw the entirety of the sill crest submerged, except for the root wads, and had waves breaking over the sill crest. In the no-sill trail, waves were breaking at the shore-line approximately six meters past what would be the sill location. The medium water with storm waves condition had breaking waves on the sill front or over the sill crest with consistent overtopping

and approximately 1-2 meters seaward of the sill location for the no-sill condition.

In the high water and average wave trial, wave breaking was mostly observed within two meters onshore of the sill rear slope with larger waves breaking on the shoreward side of the sill crest. Similar to the medium water level and average wave condition trial, the no-sill condition saw waves breaking on the shoreline. Finally, in the high water level with storm wave condition, waves were breaking within two meters shoreward of the sill crest and at the sill location for the no-sill condition.

Wave transmission and reflectionWave height attenuation through the

sill is estimated by calculating the wave transmission coefficients (Equation 9) for each sill type and hydrodynamic condi-tion (Figure 5). Flume side wall contribu-tion to wave attenuation was negligible for the scale of the experiment and therefore was excluded from the analysis. Increas-ing values of transmission coefficient illustrate larger transmitted wave heights up to a ratio value of one where the height of the transmitted wave is equivalent to the incident wave. A reflection analysis of waves returning from the sill back offshore was performed to determine the reflection coefficient for the seaward incident waves (Figure 6). Similar to the wave transmission coefficient, increasing

Figure 5 (top). Wave transmission coefficients. Wave transmission coefficient (Kt) calculated at varying breaker index values (H/h) and examined at the sill cross-shore location with average wave conditions, identified by unfilled markers, and storm wave conditions, identified by filled markers. Estimates have maximum error of Kt = 0.01.

Figure 6 (bottom). The reflection coefficient plotted vs. the wave number multiplied by water depth (kh) for storm (filled markers) and average (unfilled markers) wave conditions. Estimates have maximum error of Kr = 0.05.


Figure 7. Beach profile change pre- to post-wave event for storm wave conditions at the low water level. A) mean profile change for each sill design calculated from the spanwise average of morphology in (B-E). B)-E) plan view of the cross shore and longshore profile changes for individual sill types. Boundaries marked by dashed lines in plan view plots (B-D) illustrate each sill footprint.

values of the reflection coefficient indicate greater amounts of reflected wave energy.

To quantify the effectiveness of each sill at reducing wave heights the transmis-sion coefficient of the sill design is com-pared to that of the no-sill performance shown by blue square markers. Wave height reduction for the no-sill condition were mainly a result of wave breaking due to shoaling effects. The no-sill condition has transmission coefficients that range

from a high of 40% for the low water with storm wave conditions to a low of no wave height reduction for the medium water with average wave condition. If we compare each sill’s performance to itself, we find that the root wad sill, identified by burgundy diamond markers, show that of the six hydrodynamic conditions tested, the sill is most effective in reducing inci-dent wave heights at low water level with both wave heights. At this condition the wave height across the sill was reduced by

approximately 63% with a decreasing rate of effectiveness as the water level rises to the medium and high levels (unfilled diamond markers). For the storm wave conditions (filled diamond markers) the root wad sill transmission coefficient is again most effective at the low water level with a reduction rate of approximately 60% that decreases to 10% with rising water levels. For the rock sill, noted in Figure 5 by circular markers, the sill is most effective with average and storm

Figure 8. Beach profile change pre- to post-wave event for storm wave conditions at the medium water level. A) mean profile change for each sill design calculated from the spanwise average of morphology in (B-E). B)-E) plan view of the cross shore and longshore profile changes for individual sill types. Boundaries marked by dashed lines in plan view plots (B-D) illustrate each sill footprint. Notice elevation scale on (A) is different than in Figure 7A.


waves at low water levels, reducing wave heights by 75% and 65%, respectively. As the water level increases the effectiveness of the sill decreases equally for average and storm wave conditions to a low trans-mission coefficient value of 80% at the high water level. Lastly, the oyster hybrid sill results are shown in Figure 5 by the asterisk marker. Similar to the rock sill design, the oyster hybrid sill is the most effective at low water levels compared to other hydrodynamic conditions with average and storm wave condition wave heights across the sill are reduced by over 80% and 60%, respectively. Sill efficiency decreases with increasing water level to a low reduction rate of 30% for average waves and 20% for storm waves at high water levels compared to the no-sill condition.

A common trend of improved wave height attenuation for all sill designs is observed for all testing conditions when compared to the no-sill condition. A lin-ear trend of decreasing sill effectiveness with increasing water level is observed for all trials, which aligns with expected results as shoaling and seabed influence is reduced as water level increases. Overall the oyster hybrid sill is noticeably more effective than the root wad and rock sill designs at reducing wave heights across the sill for all water levels and wave con-ditions. There is one exception, seen with storm waves at high water level (H/h value of 0.6), where the rock sill has a greater reduction rate than the oyster hybrid sill.

The reflection coefficient (Kr), shown in Figure 6, indicates the percentage of incident wave energy that is reflected off the sill structure after impact. Calculated from the incident and reflected wave height, the coefficient is mainly influ-enced by structure material and porosity. The reflection coefficient was determined using the offshore wave gauge array where the incident and reflected wave heights were separated based on trial test-ing conditions that included the sill slope, permeability, and wavelength. A large value of Kr indicates greater amounts of reflected wave energy.

Reflection coefficients for the no-sill condition, identified by blue squares, are included in order to provide a baseline for reflection magnitude of the beach with no-sill structure. As expected, the no-sill condition has the lowest amount of reflection, attributed to the gradual

dissipative beach profile that was used for the experiment. The rock sill displays the largest reflection coefficient for almost all of the trials, as the rock sill has the lowest porosity. The root wad sill shows the over-all lowest amount of wave reflection; the low amount of reflection occurring with the root wad sill can be attributed to its porosity characteristics as the large num-ber of voids contained by the structure allows for energy to transmit through the sill. Interestingly the root wad sill has the highest reflection coefficient for the high water average wave condition (kh=1.45), potentially due to the extended height of some roots above the defined sill height, which could have resulted in larger wave interactions. The oyster sill, which had less porosity than the root wad but more porosity than the rock, was generally more reflective than the root wad but less than the rock and had a porosity in between the rock and root wad sills. All sill designs share a common trend of decreasing reflection coefficients with increased water level and shorter wave periods. The trend agrees with theory and practical knowledge as the increased water level reduces the influence the sill structure has on the wave behavior result-ing in lower reflection coefficients.

Different attempts were made in the generation of Figures 5 and 6. The wave reflection and transmission have been typically considered a linear dispersive process and, hence, a function of kh. However, under (very) shallow water con-ditions, in presence of wave breaking and nonlinear waves, and subject to the flow through porous media, it is anticipated that the wave transmission will follow a parameter representing the nonlinear character of the (non-dispersive) process (i.e. H/h), while the wave reflection is less affected by nonlinear dissipative processes and the linear behavior is still dominant. This is related to the analysis of the wave energy flux. In the energy balance equa-tion, the reflected energy flux is a function of the incident wave energy flux and the reflection coefficient, while the transmit-ted energy flux is a function of the incident energy and the dissipated energy during the transmission process. Wave energy flux as related to sill porosity is further considered in the Discussion section.

SEDIMENT TRANSPORTSill scour and sediment transportScour is an indicator of structural

instability as it attributes to the amount

of sediment erosion and undercutting occurring at the structure/sediment interface. A large amount of scour sig-nifies large amounts of wave reflection off of the sill and results in decreased horizontal and lateral support that may lead to structural failure. Scour measured for each sill for storm wave conditions at low and medium water levels are shown in Figures 7 and 8, respectively.

Scour for the low water level with storm conditions show a common trend amongst the sill designs where a scour pocket is observed at the sill front edge (Figure 7a). The magnitude is the lowest for the root wad sill (Figure 7b), followed by oyster (Figure 7d), and largest for the rock sill (Figure 7c). The scour pocket size is proportional to the amount of wave energy being reflected off the structure and turbulence generated at the structure toe. Accretion occurs at a similar distance seaward of the sill location for all de-signs, which can depend on wavelength. Minimal scour occurs on the shoreward side of each sill except for a small scour pocket observed on the shoreside of the root wad sill at a cross-shore distance of approximately 35.6 meters. Seaward accretion locations match that seen with the no-sill location, with the exception of the root wad sill where the accretion occurs closer to the sill. The root wad sill had a significant amount of scour under the sill, not quantified, which may have influenced the accretion-erosion pattern.

Scour for the medium water level with storm conditions show different mean morphological profile changes for each sill design when compared to the low water level condition (Figures 8a and 7a). The root wad sill has an increasing amount of accretion across the seaward beach profile with the highest amount of accretion, 0.1 meters, at the front of the sill (Figure 8b). Sediment erosion of quantities up to 0.3 meters were observed visually under the center of the root wad sill during post trial profile adjustment and various scour pockets behind the sill are seen in the bathymetry data. The rock sill had a mean alongshore scour depth of 5 cm at the front of the sill followed by an area of accretion and another scour pocket approximately three meters sea-ward of the sill front (Figure 8c). Minimal scouring or accretion was seen shoreward of the rock sill. The oyster hybrid sill had a small amount of accretion that extended approximately two meters seaward of the


sill followed by an area of minimal scour (Figure 8d). Like the rock, and no-sill conditions, minimal scour or accretion was observed throughout the shoreward profile of the oyster hybrid sill.

The two conditions presented were chosen because they resulted in the largest amount of scour for all tested conditions. The average storm wave conditions showed minimal scour on both the seaward and shoreward sides of each sill; with the largest amount of scour (approximately 0.02 m) measured on the seaward side of the rock sill during the low water level scenario. As the water level increased, the depth of observed scour decreased for all sill types with virtually no scour or accretion occurring at the high water average wave conditions. For the high water level with storm wave condition, a small scour pocket (0.03 m) was observed at the seaward edge of the rock sill and large scour pocket (0.06 m) was observed on the shoreward side of the root wad sill. The oyster hybrid sill most closely resembled the scour and accretion behavior of the no-sill condition, for the metric of scour, this is a positive result.

Bedload transportThe bedload sediment transport rate

for each sill is shown in Figure 9. Bedload transport is estimated using Equation (13). Data for average wave conditions at the low water level were not utilized due to ADVs being located out of the water for the trial on the onshore side of the sill. Bedload transport rates are shown at the sill seaward side (front slope) (Figure 9a), sill landward side (rear slope) (Figure 9b), front slope with reference to the no-sill condition (Figure 9c), and rear slope with reference to the no-sill condition (Figure 9d). A negative value for the bedload transport in Figure 9c and d represents a decrease in bedload sediment transport compared to the no-sill condition while a positive value indicates an increase in bedload transport compared to the no-sill condition.

Each tested sill showed an increase of bedload sediment transport compared to the no-sill condition on the seaward side of the sill (Figure 9c) and a reduction of bedload sediment transport compared to the no-sill condition on the landward side of the sill. At the sill front (Figure 9a and c) the oyster hybrid had the lowest increase of bedload transport compared to the no-sill condition followed closely by the rock sill,

Figure 9. Sediment bedload transport estimated using an energetics formulation plotted against the breaker index (H/h) at A) the sill front slope (seaward side), B) the sill rear slope (landward side). C) shows (A) removing the no sill condition loads, and D) shows (B) removing the no-sill condition loads. Storm waves filled markers and average unfilled markers.

Figure 10. Liquefaction event hydrodynamics over a 30-second window for A) the Shields parameter where sediment movement is observed by values in excess of 0.05; B) the Sleath parameter indicating plug flow formation for values greater than 0.29; and C) the pressure gradient at the sill front (seaward side), rear (landward side), and under the sill. The liquefaction event is seen in C) at approximately 14 seconds (black vertical line) shown by Dh/Dz reaching one underneath the sill (dashed line).


with both sills having an average increase in transport compared to the no-sill condition ranging from a low of no difference for the average wave conditions (H/h=0.3 to 0.4) to a high of approximately 0.05 m3/m/day for the storm wave conditions. The root wad sill had increased sediment transport on the seaward side of the sill compared to the other two designs ranging from no change in bedload transport rate at the average wave conditions, to an increase of 0.06 to

0.11 m3/m/day for storm wave conditions. In summary, all sills had increased bedload transport on the seaward side of the sill compared to the no-sill condition, with the root wad sill having the largest increase.

Results at the sill rear slope (landward side) (Figure 9b and d) follow a similar performance trend with the rock and oys-ter sills showing the greatest decrease in bedload transport rate in reference to the

no-sill condition. The rock sill had slightly better performance than the oyster sill, during storm conditions, with bedload transport reduction ranging between -0.15 to -0.3 m3/m/day compared to the no-sill condition. The root wad sill was the least effective of the three sill designs for storm wave conditions (filled markers) with bedload sediment transport reduc-tion of -0.1 m3/m/day compared to the no-sill condition. In summary, the rock

Figure 11. PDF of pore pressure gradients on sill front (seaward) slope with average wave conditions at A) low, B) medium, and C) high water levels. Pore pressure gradients on sill rear (landward) slope with average wave conditions at D) low, E) medium, and F) high water levels. Positive dh/dz values indicate decreasing soil stability with a value of positive one indicating a liquefaction event. Negative values of dh/dz indicate increasing soil stability.

Figure 12. PDF of pore pressure gradients on sill front slope (seaward) with storm wave conditions at A) low, B) medium, and C) high water levels. Pore pressure gradients on sill rear slope (landward) with storm wave conditions at D) low, E) medium, and F) high water levels. Positive dh/dz values indicate decreasing soil stability with a value of +1 indicating a liquefaction event. Negative values of dh/dz indicate increasing soil stability.


Figure 13. Wave attenuation and soil stability results for the A) root wad, B) rock, C) oyster hybrid sills at i) low, ii) medium, and iii) high water levels with storm wave conditions. Arrows and wave features represent the wave transmission and wave height changes across the sill, respectively. Wave features and arrows that are larger signify high wave heights, transmission, and reflection coefficients. Scour is shown through drawn erosion features in the beach profile and liquefaction potential is represented by the proximity of black dots, signifying sediment grains, beneath the sill. Dots closer in proximity detail a stable bed with low liquefaction potential, while dots with greater distance represent an increasingly unstable soil with higher liquefaction potential.and oyster sills were more effective at re-ducing scour on the shoreward side of the sill than the root wad sill, yet all sill types were more effective at erosion reduction compared to the no-sill condition.

Sediment transport statistics for both the seaward and landward sides of the sill are consistent with observations of scour shown in Figures 7 and 8. All sills had more sediment transport and scour on the seaward side of the sill when compared with the no-sill condition, and all sills had less sediment transport and scour on the landward side compared with the no-sill condition. Additionally, scour and sedi-ment transport estimates show that the root wad was most vulnerable to higher rates of sediment transport on both sides of the sill compared to the rock and oyster hybrid sills. However, with scour measure-ments, we do see that the rock sill had consistently more scour on the seaward side of the sill than all other sill types.

Liquefaction potentialSill stability, in relation to sediment

transport, not only relies on shear driven transport (bedload sediment transport and scour), but also is influenced by

momentary liquefaction of the sill un-derlying sediments. Figure 10 shows what a momentary liquefaction event looks like in time. Figure 11 and 12 then present probability distributions of pore water pressure gradients for sill front and rear slopes for average and storm wave conditions to demonstrate the percent-age of time that each sill was exposed to momentary liquefaction conditions over the 44-minute trials. The pressure gradient for each sill type was examined for liquefaction potential identified by a positive dh/dz value equal to or greater than one. Narrow banded results illustrate higher soil stability and low liquefaction potential attributed to minimal pressure gradient fluctuations. Results that follow a broader banded trend show higher liq-uefaction potential and soil instabilities as the pressure gradient fluctuations are greater in magnitude and detail the large oscillating pressure forces acting on sedi-ment grains. Negative pressure gradient values indicate positive stabilizing pres-sure due to the downward seepage. The downward flow of water acts in the same direction of gravity resulting in increased soil stability. Positive pressure gradient

values show time events of upward wa-ter flow and decreased soil stability as a result of the water pressure acting in the opposite direction of gravity.

Pore water pressure gradients at the front slope, rear slope, and underneath the rock sill at high water level and storm wave conditions are seen over a 30 second time window in Figure 10c. A liquefaction event is shown at a time of 14 seconds when the vertical hydraulic head gradient underneath the rock sill reaches a value of one. The Shields parameter, Figure 10a, and the Sleath parameter, Figure 10b, were calculated to quantify the movement of sediment at the bed-level. A Shields pa-rameter of 0.05 indicates the sediment has surpassed the threshold for motion (sheet flow sediment transport occurs at a Shields parameter of 0.8), while a Sleath parameter of 0.29 indicates initiation of plug flow formation (Shields 1936; Sleath 1999).

The Shields parameter at the sill front and rear are both greater than the threshold for motion at the time of the liquefaction event (Figure 10a). The Sleath values at the rock sill front and rear slopes are less than the limiting threshold


Figure 14. Sill surface area to volume ratio plotted against wave energy dissipation for a) average (unfilled markers) and b) storm wave conditions (filled markers) normalized against the no-sill condition. The root wad, rock, and oyster sills are identified by SA/Vsill values of 1.6, 1.9, and 3.6, respectively. Triangular, circular, and square markers identify water levels at the sill of 0.33, 0.61, and 0.91 meters respectively.

value signifying the absence of sediment plug flow on either side of the sill (Figure 10b). From the vertical hydraulic head gradient (Figure 10c), we see that there was liquefaction of the sediment occur-ring under the sill, but not to either side at a time of 14 seconds. Results show that moments of increased shear (indicated by the Shields parameter) and pressure driven sediment transport as indicated by Sleath and the vertical hydraulic head gradient at the front of the sill, bearing the brunt of wave impact, may be indica-tive of potential momentary liquefaction events underneath of the sill.

A reduction in pressure gradients for all sill types is seen as the water levels increase with the exception of the rock sill at the high water level with average wave conditions (Figure 11f). The observed decreasing trend follows theoretical ap-plications as bed level orbital velocities induced by small amplitude waves have lessening impacts on the sea floor as water depth increases. The no-sill condition, shown by the dotted line, is used as the baseline for soil stability and liquefaction potential. Best observed in the sill front slope pressure gradients (Figure 11a, b, c), pressure gradient fluctuations for the no-sill condition (dh/dz) become more narrow with increasing water depth. This indicates a stable soil with very low liquefaction potential.

At the low water level (Figure 11a) that the oyster hybrid sill is the most unstable with a (dh/dz) range of -0.3 to +0.3, which despite being the most unstable, still is representative of a low liquefac-tion potential as the (dh/dz) fluctuations are small in scale. An opposite trend is seen for the rear slope pressure gradients where at the lowest water level (Figure 11d) the oyster hybrid sill is the most stable and the rock sill has the greatest pressure gradient fluctuation range of -0.15 to +0.15. As the water level increases to medium (Figure 11b, e) and high (Figure 11c, f) the oyster hybrid and root wad sills become more narrow banded indicating increased stability while the rock sill distribution flattens indicating decreased soil stability. All sills have a broader range of pressure gradients at the sill front slope (Figure 11a, b, c) when compared to the no-sill condition. The difference could be a result of surface waves impacting the sill structures in combination with the additional normal stress due to sill weight. Conversely, with

the exception of the rock sill, the distribu-tion the oyster hybrid and root wad sills closely resemble the behavior of the no-sill condition on the sill rear slope (Figure 11d, e, f) indicating minimal changes to sediment stability. This could be due to minimal wave impact on the rear slope of the sill due to reduced wave heights. The rock sill deviation at the medium (Figure 11e) and high (Figure 11f) water conditions may be due to transmitted wave energy from overtopping events. No liquefaction events were observed for any of the sill designs with the average wave testing conditions.

A noticeable difference is seen be-tween the pressure gradients for average (Figure 11) and storm wave (Figure 12) conditions, attributed to the storm sur-face gravity waves creating greater pres-sure forces on the sill and underlying sed-iments. Gradients for the sill front slope show large deviations when compared to rear slope results for the low and medium water levels (Figure 12a and b compared

to d and e), but have closer trends at the high water level (Figure 12c and f). Simi-lar to the results shown in Figure 11, the oyster hybrid sill was the most unstable at the low water level (Figure 12a) where pressure gradient fluctuations ranging from -0.6 to +0.6 are seen. Like with the average wave conditions, the stability of the oyster hybrid sill improves with increasing water level as the fluctuations are skewed yet truncated (Figure 12a, b, c). The rock sill, identified by the solid line, shows the most stability in the low water level (Figure 12a), but then quickly losses stability as the water level increases leading to a liquefaction and near lique-faction events that are seen at the high and medium water levels, respectively (Figure 12c, b). On the rear slope of the sill the oyster hybrid and root wad sills have closely related performances with the root wad sill showing more stability at the medium water level (Figure 12e). The rock sill again shows decreasing stability with increasing water levels (Figure 12d, e, f). Gradients for the sill rear slope are narrower but still have a major percentage of positive pressure gradients indicat-ing soil instability for all three sills. The observed broad gradients throughout Figure 12 detail large changes between downward and uplift pore pressure forces acting on the sediment.

With the exception of the front slope low water condition (Figure 12a) and the rear slope medium water level (Figure 12e), the oyster hybrid sill results are the most similar to the no-sill condi-tion in terms of both soil stability and liquefaction potential, where in general, the no-sill case has the least amount of liquefaction and near liquefaction events. For the purpose of sill stability, liquefac-tion could cause the sill to settle and or fail depending on the magnitude of the event and reoccurrence frequency. A sill with a lower height has implications to attenuate less wave energy.

DISCUSSION Sill performance

Figure 13 is a cartoon that summarizes experimental results for storm wave con-ditions at low, medium, and high water levels with storm wave conditions. The figure represents how wave height, wave transmission, wave reflection, sill scour, and bed momentary liquefaction change for each hydrodynamic condition over the rock, oyster hybrid, and root wad sills.


Overall the oyster hybrid sill was the most effective for the storm wave condi-tions at all water levels. Our observa-tions show that the oyster hybrid sill was very effective at reducing incident wave heights, we believe its effectiveness is due to its designed use of transmission through the small pore spaces between the oyster shell for energy dissipation, whereas the rock sill uses wave reflection for dissipation, which is not as efficient. We observed that the root wad sill also uses transmission as a means of wave energy dissipation; however, unlike the oyster hybrid’s porous design that con-tains a significant amount surface area due to the shell, the root wads contain a lot of open space, allowing the waves to maintain energy as they pass through. To quantify these observations, Figure 14 shows a measure of the wave energy dissipation across each sill plotted against a measure of the sill porosity and volume. Where the wave energy dissipation (ε) across the sill (Δx) is estimated using the wave energy flux per cross shore distance traveled (see Equation (22)), where, E is the wave energy estimated using wave height, and cg is the wave group velocity. We estimated the effect of combined sill porosity and mass based on sill material and geometry characteristics using the sill material total surface area per sill unit volume estimate (SA/Vsill), SA and Vsill are described in the methods section. Figure 14 shows that the root wad sill has the smallest SA/Vsill compared to the other two sills, the root wad also had the low-est amount of energy dissipation across the sill for both wave conditions (aver-age 14a, storm 14b) and each water level (markers: triangle – low, circle – medium, square – high). Although the rock sill is characteristically different than the root wad sill, it had similar SA/Vsill due to its low porosity and large mass. Finally, the oyster sill had the largest SA/Vsill. During the average wave conditions (Figure 14a) the oyster sill had the most wave energy dissipation for each water level tested, more than both the rock and the root wad sills. During the storm wave conditions (Figure 14b) at the low water level the oys-ter sill had more wave energy dissipation than the rock sill, but at the high wave energy conditions the oyster and rock sills performed the same. Although SA/Vsill is shown as a function of wave energy dis-sipation, the SA/Vsill metric shows similar trends with scour and sediment transport quantities. Overall, results show that sill

porosity is important to both its ability to dissipate incident wave energy as well as reduce scour and liquefaction potential.

For liquefaction potential, all the sill designs were relatively stable at all water levels with average wave conditions, but showed varying stability behavior with storm wave conditions. The stability per-formance of the oyster sill with the storm wave conditions can be relaed to its light weight and wave attenuation capability. The root wad sill, while lighter than the oyster hybrid sill, was not very effective at reducing incident wave heights, which resulted in larger wave pressure acting on the sill foundation contributing to higher observed pressure gradients. Like the oyster hybrid sill, the rock sill was very effective at reducing incident wave heights, but the sill used more rock, which has a substantially higher unit weight compared to oyster shell, and again did not reduce wave energy in the same way (rock sill uses reflection instead of dis-sipation of wave energy), resulting in greater excess normal stress applied on underlying sediment particles.

Lastly, scour depths for each sill can be related to the sill’s method and rate of wave energy dissipation. The rock sill design uses wave reflection as its energy dissipation method, which relies on re-directing the energy vector verses dis-sipating the energy. The reflected energy causes a lot of turbulence and changes the velocity field at the redirection interface, resulting in higher sediment mobilizing forces and scour, and explains why large scour pockets are observed at the seaward (front) edge of the rock slope. The root wads use transmission for wave energy dissipation similar to the oyster hybrid; however, its lessened ability to signifi-cantly reduce wave energy coupled with the lack of a foundational geotechnical liner resulted in large amounts of scour underneath and behind the sill and a bar like accretion area at the seaward side of the sill. The oyster hybrid sills transmis-sive properties combined with its light-weight construction and large amount of surface area due to the shell characteristics enabled it to effectively attenuate the waves without significantly impacting soil stability. Additionally, large amounts of energy dissipation due to the oyster hybrid sills transmissive property reduced the amount of foundational scour around the sill, which will result in greater struc-tural integrity over longer time periods.

Damage to sillsDespite a short testing time of only 44

minutes, damage of varying magnitudes was observed in all three sill designs during the storm wave condition trials. The greatest amount of damage was seen with the root wad sill as the large amount of wave turbulence occurring at the sill location caused significant erosion of the sediment underneath the sill. This resulted in a settlement of approximately eight cm causing the anchoring system to slacken, and increasing the displacement of the sill due to wave impact. One past experiment presented at a 2014 ASBPA conference showed that long-term use of large woody debris in coastal areas may lead to instability of the structure, but further investigation should be per-formed and published. Throughout test-ing the oyster sill settled approximately 15 cm and multiple bags were displaced vertically but were not dislodged from the structure due to the overlaying poultry wire. Minimal damage was seen with the rock sill, however; with longer testing periods a greater level of damage would be expected as large scour pockets were observed during storm trials, and the sill was most vulnerable to liquefaction of sediments underneath. Over the duration of a full scale storm, scour pocket growth and liquefaction could lead to the collapse of the sill frontal wall or reduction in wave attenuation ability due to settlement.

CONCLUSIONS Our experiment used wave conditions

estimated from historical wind data and bay lengths to examine the wave attenu-ation capabilities and soil stability prop-erties of three different living shoreline designs (rock, oyster hybrid, root wad) over two wave and three water level con-ditions. Wave and water level conditions as well as sill design heights are placed into the context of and applicable to field observations on the east, west, and Gulf coasts of the U.S. All sill designs were found to have better wave attenu-ation capabilities and varying sediment stability thresholds when compared to a no-sill condition. Over all, we found the oyster hybrid sill comprised with a rock core to be the most effective in reducing wave energy and wave heights reaching the shoreline and performed the best with respect to sediment transport vulner-ability, mirroring soil stability proper-ties of the no-sill condition. While sill performance is an important aspect in


deciding the best design for a proposed location, other important factors include ecological benefits, material availability, construction costs, and storm frequency. Each factor must be considered during design and will be different depending on geographic location. Depending on site location, it may be more feasible to use a different material for design. Bilkovic et al. (2017) weighs various attributions of living shoreline design from perspec-tives of managers, engineers, ecologists, and biologists. It may be that a root wad sill is cheaper and sources local material for the site, or that a rock sill is the most practical option in an area that could not support an oyster population. Every site is unique, and every coast around the U.S. has eroding marshes and is in some state of implementing marsh restoration techniques to cope with future storms and water levels. Results from our ex-periments are meant to aid in discussion for sill design alternatives within site constraints for living shoreline marsh restoration projects.

Overall, lessons learned using the three sill designs can be extended beyond sill material. The experiment showed the importance of sill porosity and weight to both sill stability and sill wave energy re-duction capability. We saw that sills with low (rock sill) and high (root wad sill) porosity resulted in higher levels of en-ergy reflection vs transmission and high sediment transport at the sill location, leading to scour. Sills with more mate-rial surface area for energy dissipation (oyster) encouraged increased energy

dissipation across the sill when compared to materials with less transmissive sur-face area (rock). Sills with high material weight (rock) increased pressure loading on the sediment resulting in decreased soil stability and a higher liquefaction po-tential during storms. Lastly, results from the experiment provide an opportunity for expansion into different sill material categories including the possible use of synthetic sill designs, where our results could aid in design of synthetic sill geom-etries and material characteristics. Future work could expand into testing synthetic designs meant to mirror the lightweight and complex structure of root wads with increased surface area of oyster shell to improve transmissive dissipation.

ACKNOWLEDGEMENTS We thank the Graduate Women In

Science for providing funding for our ex-periment. We thank the National Science Foundation (NSF) and Natural Hazards Engineering Research Infrastructure (NHERI) for providing the instrumen-tation and sediment for the experiment as part of NSF Award# 1756449 and the Oregon State University College of Engineering for providing flume time. We thank the U.S. Coast Guard for sup-porting JC. We also thank Oregon State University faculty members including Matt Evans and Tim Maddux for their assistance with geotechnical engineer-ing aspects and wavelab equipment operation. Lastly, we thank Tyler Messie, Seok-Bong Lee, Hailey Bond, Elizabeth Holzenthal, and Jeremy Smith for their help with sill construction and testing.

REFERENCES Allen, R.J., and B.M. Webb, 2011. “Determination

of wave transmission coefficients for oyster shell bag breakwaters.” In Coastal Engineer-ing Practice — Proc. 2011 Conference on Coastal Engineering Practice. https://doi.org/10.1061/41190(422)57.

Beseres Pollack, J., Yoskowitz, D., Kim, H. and P.A. Montagna, 2013. “Role and value of nitrogen regulation provided by oysters (Crassostrea Virginica) in the Mission-Aransas Estuary, Texas, USA.” PLoS ONE. https://doi.org/10.1371/journal.pone.0065314.

Bilkovic, D.M., and M.M. Mitchell, 2017. “Design-ing living shoreline salt marsh ecosystems to promote coastal resilience.” In Living Shorelines: The Science and Management of Nature-Based Coastal Protection, CRC Marine Science, 293-316.

Bilkovic, D.M., Mitchell, M., Mason, P., and K. Duhring, 2016. “The role of living shorelines as estuarine habitat conservation strategies.” Coastal Management. https://doi.org/10.1080/08920753.2016.1160201.

Bouws, E., Günther, H., Rosenthal, W., and C.L. Vin-cent, 1985. “Similarity of the wind wave spec-trum in finite depth water: 1. Spectral form.” J. Geophysical Research, 90(C1), 975-986.

Dall, P.C., 1979. “A sampling technique for littoral stone dwelling organisms.” Oikos. https://doi.org/10.2307/3544518.

Davis, J.L.D., 2017. “Gaps in knowledge information we still need to know about living shoreline erosion control.” In Living Shorelines: The Sci-ence and Management of Nature-Based Coastal Protection, Michael Kennish and Judith Weis, eds., 1st ed., Boca Raton: CRC Marine Sci-ence, 481-503.

Dean, R.G., and R.A. Dalrymple, 1984. “Engi-neering wave properties.” In Water Wave Mechanics for Engineers and Scientists., 2nd ed., 78-130. Hackensack: World Scientific. https://doi.org/10.1029/eo066i024p00490-06.

Faherty, M., 2011. “Oyster reef restoration and monitoring, Wellfleet, MA Draft Final Re-port.” South Wellfleet.

Foster, D.L., Bowen, A.J., Holman, R.A. and P. Natoo, 2006. “Field evidence of pres-sure gradient induced incipient motion.” J. Geophysical Research: Oceans. https://doi.org/10.1029/2004JC002863.

Gittman, R.K., Scyphers, S.B., Smith, C.S., Neylan, I.P, and J.H. Grabowski, 2016. “Ecological consequences of shoreline hardening: a meta-analysis.” BioScience. https://doi.org/10.1093/biosci/biw091.

Griggs, G., 2005. “The impacts of coastal armoring.” Shore & Beach, 73 (1), 13-22.

Hardaway, C.S, and R.J. Byrne, 1999. “Shoreline management in Chesapeake Bay.” Special Report in Applied Marine Science and Ocean Engineering Number 356. Virginia Sea Grant Publication VSG-99-11.

Hardaway, C.S., Milligan, D.A., and C.A .Wilcox, 2017. “Living shoreline design guidelines for shore protection in Virginia’s estuarine environments.” Virginia Institute of Marine Science 2 (Special Report in Applied Marine Science and Ocean Engineering #463).

Harman, W.A., Jennings, G.D., Tweedy, K.R., Buck, J.A. and D.L. Taylor, 2001. “Lessons learned from designing and constructing in-stream structures.” In Proc. 2001 Wetlands Engineer-ing and River Restoration Conference. https://


doi.org/10.1061/40581(2001)71.Henderson, J., and J. O’Neil, 2003. “Economic

values associated with construction of oyster reefs by the Corps of Engineers.” EMRRP Technical Notes.

Holtz, R.D., and W.D. Kovacs, 1981. “Water in soils, II: permeability, seepage, effective stress.” In An Introduction to Geotechnical Engineering, 1st ed., Englewood Cliffs: Prentice-Hall, 199-282.

Hsu, T.J., Elgar, S. and R.T. Guza, 2006. “Wave-in-duced sediment transport and onshore sand-bar migration.” Coastal Engineering. https://doi.org/10.1016/j.coastaleng.2006.04.003.

Hudson, R.Y., 1974. “Concrete armor units for protection against wave attack. Report of ad hoc committee on artificial armor units for coastal structures.” U.S. Waterways Experi-ment Station miscellaneous paper.

Lomonaco, P., Vidal, C., Losada, I.J., Garcia, N., and J.L. Lara, 2005. “Flow measurements and nu-merical simulation on low-crested structures for coastal protection.” In Environmentally Friendly Coastal Protection. Springer, Dor-drecht, 191-210.

Kuykendall, K.M., Moreno, P., Powell, E.N., Soniat, T.M., Colley, S., Mann, R., and D.M. Munroe, 2015. “The exposed surface area to volume ratio: is shell more efficient than limestone in promoting oyster recruitment?” J. Shellfish Re-search. https://doi.org/10.2983/035.034.0203.

Miller, J.K., Rella, A., Williams, A., and E. Sproule. 2015. “Living shorelines engineering guide-lines.” Stevens Institute of Technology.

Mitchell, M., and D.M. Bilkovic. 2019. “Embracing dynamic design for climate-resilient living shorelines.” J. Applied Ecology. https://doi.org/10.1111/1365-2664.13371.

Moody, J.A., Gentry, M.J., Bouboulis, S.A., and D.A. Kreeger, 2020. “Effects of substrate (protec-tion and type) on Ribbed Mussel (Geukensia Demissa) recruitment for living shoreline applications.” J. Coastal Research. https://doi.org/10.2112/jcoastres-d-19-00062.1.

Nielsen, P., 1992, “Sediment mobility, bed-load and sheet-flow.” In Coastal Bottom Bound-ary Layers and Sediment Transport, 4th ed., Hackensack: World Scientific, 95-128.

Panofsky, H.A., and Dutton, J.A., 1984. Atmospheric Turbulence: Models and Methods for Engineer-ing Applications, Wiley Press, New York, NY, 1st ed. 389p.

Priest III, W.I. 2017. “Practical living shorelines

tailored to fit in Chesapeake Bay.” In Living Shorelines: The Science and Management of Nature-Based Coastal Protection, edited by D.M. Bilkovic, CRC Marine Science, 185-210.

Rafferty, M., 2017. “Computational design tool for evaluating the stability of large wood structures.” USDA Technical Note TN-103.2.

Schlicker, H.G., Gray, J.J., and Bela, J.L., 1978. “Rock material resources of Benton County, Oregon.” Oregon Department of Geology and Mineral Industries Bulletin 27, 49p.

Scyphers, S.B., Powers, S.P., Heck, K.L, .and D. By-ron, 2011. “Oyster reefs as natural breakwaters mitigate shoreline loss and facilitate fisheries.” PLoS ONE. https://doi.org/10.1371/journal.pone.0022396.

Seal, R., Stein, O.R., and S.F. Boelman, 1998. “Perfor-mance of in-stream habitat structures under flood conditions.” In Proc. ASCE Wetlands Engineering River Restoration Conference.

Shields, A., 1936. “Application of similarity prin-ciples and turbulence research to bed-load movement.” Translated by W.P. Ott and J.C. Uchelen. Soil Conservation Service.

Shih, P.K., and W.L. Chang. 2015. “The effect of water purification by oyster shell contact bed.” Ecological Engineering. https://doi.org/10.1016/j.ecoleng.2015.01.014.

Sleath, J.F.A., 1999. “Conditions for plug forma-tion in oscillatory flow.” Continental Shelf Research. https://doi.org/10.1016/S0278-4343(98)00096-X.

Sumer, B., and Fredsøe, J., 2002. “Impact of liquefac-tion” In The Mechanics of Scour in the Marine Environment, World Scientific, Singapore, 445-520, https://doi.org/10.1142/4942.

Swann, L.. 2008. “The use of living shorelines to mitigate the effects of storm events on Dauphin Island, Alabama, USA.” American Fisheries Society Symposium.

Sylte, T.L., and J.C. Fischenich, 2000. “Rootwad composites for streambank stabilization and habitat enhancement.” EMRRP Technical Notes Collection (ERDC TN-EMRRP-SR-21), U.S. Army Engineer Research and Develop-ment Center, Vicksburg, MS.

Szabados, M.W., 1982. “Intercomparison of the off-shore wave measurements during ARSLOE.”

Terzaghi, K. 1943. “Stress conditions for failure in soils.” In Theoretical Soil Mechanics, 8th ed., New York: John Wiley and Sons, 7-25.

U.S. Army Corps of Engineers, 2011. “Part VI: Chapter 5 Fundamentals of Design.” Coastal

Engineering Manual.U.S. Army Corps of Engineers, 2015. “Part II:

Chapter 2 — Meteorology and Wave Climate.” Coastal Engineering Manual.

van der Meer, J.W., and I.F.R. Daemen, 1994. “Stability and wave transmission at low-crested rubble-mound structures.” J. Wa-terway, Port, Coastal and Ocean Engineer-ing. https://doi.org/10.1061/(ASCE)0733-950X(1994)120:1(1).

Vidal, C., Losada, M.A., Medina, R., Mansard, E.P.D., and G. Gomez-Pina, 1993. “A universal analysis for the stability of both low-crested and submerged breakwaters.” In Coastal En-gineering 1992, 1679-1692.

VIMS CCRM, 2017. “Living shorelines: design op-tions – marsh sill with planted marsh.” VIMS Coastal Resource Management Center. http://ccrm.vims.edu/livingshorelines/design_op-tions/marsh_sill_planted.html.

Vincent, C.L., 1985. “Depth-controlled wave height.” J. Waterway Port Coastal and Ocean Engineering-ASCE, 111(3), 459-475.

Walker, R., Bendell, B., and L. Wallendorf, 2011. “Defining engineering guidance for living shoreline projects.” In Coastal Engineer-ing Practice — Proc. 2011 Conference on Coastal Engineering Practice. https://doi.org/10.1061/41190(422)86.

Wengrove, M.E., Foster, D.L., Lippmann, T.C., de Schipper, M.A., and J. Calantoni, 2019. “Observations of bedform migration and bedload sediment transport in combined wave-current flows.” J. Geophysical Re-search: Oceans, 124(7), 4572-90. https://doi.org/10.1029/2018JC014555.

Whitman, E.R., and M.A. Reidenbach, 2012. “Ben-thic flow environments affect recruitment of Crassostrea Virginica larvae to an intertidal oyster reef.” Marine Ecology Progress Series. https://doi.org/10.3354/meps09882.

Zelt, J. and Skjelbreia, J. E., 1992. “Estimating incident and reflected wave fields using an arbitrary number of wave gauges.” Coastal Engineering Proc., 1(23), 777-789. https://doi.org/10.9753/icce.v23.%25p.

Zhu, L., Chen, Q., Wang, H., Capurso, W., Nie-moczynski, L., Hu, K., and G. Snedden. 2020. “Field observations of wind waves in upper Delaware Bay with living shorelines.” Estuaries and Coasts. https://doi.org/10.1007/s12237-019-00670-7.

Observations of wave attenuation, scour, and subsurface...

Documents

Transcript of Observations of wave attenuation, scour, and subsurface...