# Numerical Simulations of the Hydraulic Performance of a Breakwater-Integrated Overtopping Wave Energy Converter

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## Abstract

**:**

## 1. Introduction

^{2}) that is greater (wind: 0.4–0.6 kW/m

^{2}; solar: 0.1–0.2 kW/m

^{2}) [12]. However, the marine technologies, the Wave Energy Converters (WECs) among others, are still at immature stages to be commercialized successfully in the market [13,14]. Despite this, more than 1000 WECs have been developed worldwide and more than 140 are patented [15].

_{r}and the average overtopping flow inside the reservoir q

_{reservoir}; (2) the structural performance in terms of pressures acting on the device; and (3) the harbor safety by measuring the average overtopping flow at the rear side of the rear wall q

_{rear}. Section 4 discusses the results obtained with reference to the OBREC pilot plant installed in the port of Naples in 2017. Some conclusions and recommendations on the OBREC design optimizations are drawn in Section 5.

## 2. Material and Methods

#### 2.1. The Overtopping Breakwater for the Energy Conversion Cross Section

_{r}is the reservoir width, B

_{s}the horizontal extension of the sloping plate, α its inclination (equals to 34° in the physical model), d

_{w}is the height of the sloping plate and R

_{r}and R

_{c}are the freeboard crests (with respect to the mean sea water level) of the sloping plate and of the rear wall, respectively.

_{w,low}= 2.25 m and d

_{w,high}= 3.75 m at prototype scale. In this paper, the numerical model is calibrated considering only the d

_{w,high}configuration.

_{n50}= 40 mm for the armour layer, D

_{n50}= 20 mm for the filter layer and D

_{n50}= 2 mm for the core. In the numerical model, the aforementioned porous layers are characterized by several parameters as the porosity itself n, the added mass coefficient c

_{A}, the linear α and non-linear β friction coefficients. These last two parameters describe respectively the laminar and the turbulent flow properties in between the stones by means of the Forchheimer equation [35], defined as

_{f}is the porous flow velocity averaged over the total sample area (grains), g the gravitational acceleration and υ the fluid kinematic viscosity. In Equation (1), the third term on the right-hand side is usually small and therefore neglected. The β coefficient has been set equal to 1000, while α was set to 1.1, 1.0 and 0.8 for the armour, the filter and the core layers, respectively. These values have been defined based on the literature [36,37,38].

#### 2.2. Wave Conditions and Measurements

_{m0}, the peak frequency f

_{p}and the so-called peak enhancement factor γ (3.3 in all the tests). The wave series considered at least 1000 waves.

_{m0}, γ, the spectral wave period T

_{m−1,0}, the water depth in the generation area h and the duration of the simulation t. The simulation program was divided into two test series, “normal” (representing poor and mild wave climate) and “extreme” (representative of a severe storm condition). Table 1 contains the characteristics of both the normal and extreme wave conditions. Each numerical test considers at least 500 waves, which are sufficient to perform a statistical wave overtopping analysis as demonstrated by Romano [39].

_{m0}(0.077 m)—are characterized by the same H

_{m0}but different spectral wave periods T

_{m−1,0}. Therefore, the tests chosen allowed the analysis of the hydraulic performance according to the wave steepness s

_{o}.

- the reflection coefficient K
_{r}, in normal conditions; - the average overtopping discharge inside the reservoir q
_{reservoir}, in normal conditions; - the pressures p acting on the structure, under extreme conditions;
- the average overtopping discharge at the rear side of the crown wall q
_{rear}, under extreme conditions.

_{r}, by using WG1, WG2 and WG3 in Figure 3, located in front of the structure. The offshore wave gauge, WG0, has been used to measure the generated wave height.

_{rear}) was determined in the lab by using a ramp to guide the overtopping wave volumes into a box. A depth gauge was installed in the box to measure the overtopping discharge and to control the pump to empty the box at a given threshold level. Similarly, the overtopping discharge in the front reservoir (q

_{reservoir}) was measured using depth gauges, which controlled several pumps. The 2-D model has been provided with a pipe to let the overtopping water flows towards the inshore edge, and two WGs were placed on the top of the sloping plate and at the rear side of the rear wall to estimate respectively the values of q

_{reservoir}and q

_{rear}(Figure 4). The overtopping discharges have been derived by integrating (along the vertical) cell by cell the horizontal velocity component multiplied by the cell height (i.e., z direction).

#### 2.3. Numerical Wave Flume and Mesh

#### 2.4. Calibration of the 2-D Numerical Model

_{r}and q

_{reservoir}under normal conditions. This procedure implied the variation of the porosity n assigned to the layers of the breakwater, under the same wave attack. Table 2 reports the values assigned to each layer together with the comparison between the experimental q

_{res,exp}and the numerical q

_{res,num}discharge flows. The final n has been defined according to the best agreement, i.e., configuration 2 in Table 2.

_{r}, q

_{reservoir}and therefore the reliability of the numerical model. In order to compare the numerical q

_{reservoir}with the experimental values and the theoretical formulations, they have been non-dimensionalised. Furthermore, by following the definition proposed by Ahrens and Heimbaugh (1986, Equation (4)) [42], it has been possible to take into account the difference between the wave spectra generated in the lab and in the numerical wave channel, thanks to the presence of the significant wave height H

_{m0}in Equation (4). The theoretical formulae considered to evaluate the overtopping discharge are reported in Equations (5) and (6) and correspond to those adopted in the EurOtop 2016 [43] to assess the same phenomenon on the slopes such as dikes, levees and embankments.

_{off}is the offshore slope angle, while γ

_{b}is the berm influence factor, γ

_{f}is the roughness influence factor, γ

_{β}is the oblique wave action influence factor, γ

_{ν}is the influence factor for a vertical wall, ξ

_{m−1,0}is the breaker parameter and R

_{r}is the freeboard crest of the sloping plate.

_{o}. The reason can be addressed to the slight modification occurred in modelling the cross section for the numerical analysis, concerning the thickness of the impermeable part as described in Section 2.1. (Figure 2b,c). However, such modification has no implication in the representation of the discharge flows inside the reservoir, i.e., q

_{reservoir}. In fact, Figure 6 shows that the numerical q*

_{reservoir}gives a better estimation of the laboratory results. This is confirmed by the performed statistical analysis. Indeed, the difference among the experimental, the numerical and the theoretical values obtained are statistically computed by means of 2 parameters, describing the error made, i.e., the root mean square error (RMSE) and the Willmott index (WI) [44], defined in Equations (7) and (8), respectively. The quantity “y” represents the dimensionless discharge rate, defined in Equation (4). The subscripts “s” stands for the experimental data, while “mod” for the numerical and the theoretical ones are according to the analysis performed. Therefore, “ӯ

_{s}” is the mean of the experimental discharge flows. A good representation of the experimental data is characterized by a RMSE value close to 0 and WI values close to 1. Indeed, the numerical RMSE and WI are equal to 0.005 and 0.75, while the theoretical ones to 0.007 and 0.59, respectively.

#### 2.5. Tested Configurations

_{r}and the shape of the sloping plate (its inclination a), while the sloping plate freeboard R

_{r}and its longitudinal dimension of B

_{s}have been kept constant. Figure 7 shows all the tested configurations (named M1–M5), which the main parameters of are synthesized in Table 3.

_{r}. The two cross sections imply a variation of B

_{r}of 25% with respect to the benchmark case. The aim of this analysis is to assess its relevance in the saturation of the reservoir, the maximization of the energy production, the limitation of the pressures along the crown wall and the values of q

_{rear}.

## 3. Results

_{r}, q

_{reservoir}, q

_{rear}and pressures are used to improve the design of the device, by providing some general guidelines according to the analysis of the wave–structure interactions.

#### 3.1. Effects of Geometric Changes on the Hydraulic Performance

_{r}and non-dimensional overtopping at the rear of the structure, q*

_{reservoir}, under normal conditions.

_{r}versus the steepness s

_{o}for all the configurations, under normal conditions. The laboratory results (already shown in Figure 5) is respected. The configuration M2 leads to greater values with respect to M1. This result was expected considering the work of Zanuttigh [47], who demonstrated that structures characterized by a submerged berm lead to a smaller K

_{r}than those with a straight slope validated also by the study performed by Formentin [48]. The design of a toe protection in case of an OBREC installation in a breakwater without berm should be taken into account.

_{r}does not significantly affect the wave reflection as it can be derived from the similar values of K

_{r}for M2, M3 and M4. The value of B

_{r}can therefore be selected based on the maximization of the energy production and on the minimization of the costs, while adapting the overall dimension of the device to the spatial constraints posed by the installation in existing breakwaters.

_{m0}. As a matter of fact, the presence of a vertical element in the ramp shape has a greater influence on the smallest wave height. Therefore, it is important to consider the response of the structure with respect to the wave condition characterized by the greatest frequency in a typical wave climate.

_{reservoir}for M1–M5 are compared with the results related to M2. The absence of the berm does not change the values of q*

_{reservoir}with respect to M1, at least for the tests characterized by the greatest discharge rates and so the greater wave heights. Therefore, the higher reflection does not change in a significant way the potential power production.

_{reservoir}for M3 suggest that the reservoir is under-dimensioned. The greater reservoir size of M4 does not lead to a real improvement of the discharge rate. Therefore, it can be concluded that this parameter does not affect significantly the hydraulic performance of the device. Therefore, its design in existing breakwaters has to be focused on the height and the shape of the sloping plate that directly influence the overtopping phenomenon. The dynamic inside the reservoir is more connected to the position of the pipes, which lead the water to flow towards the turbines.

#### 3.2. Effects of Geometric Changes on the Structure Loads

_{r}. In M5 the sensors related to the sloping plate were shifted along the orthogonal direction that links the original inclination to the new ones.

_{250}, which corresponds to the non-exceedance level of about 99.7%. Table 4 reports the values related to the sloping plate, the crown wall and the bottom part of the reservoir (uplift pressures).

_{250}similar to M1.

_{250}are reported in Table 5. All the configurations show results very similar to each other, being the inshore corner, i.e. corresponding to the gauge 10

_{in}(see Figure 4), the most stressed part of the reservoir. This is not completely true for the configurations M4 and M5 due to the size of B

_{r}and to the lower number of overtopping waves, respectively.

#### 3.3. Harbor Safety

_{rear}, i.e., the average overtopping discharge at the rear side of the crown wall, in extreme conditions (see Table 1).

_{rear}measured for the OBREC case (Figure 2b) were higher compared to a traditional rubble mound breakwater with similar overall dimensions (Figure 2a) [30]. Therefore, a parapet (nose) has been introduced on the top of the crown wall to increase the safety level of the area inshore the structure. The role of the parapet is to redirect the up-rushing waves back into the front reservoir and towards the sea.

_{rear}, such as the crown wall height h

_{wall}, the extension of the nose h

_{n}and the inclination of the parapet ε (Figure 10).

_{rear}. The optimal range for ε was found between 30° and 45°, which combines a good reduction of q

_{rear}and a limited increase of the pressures on the wall in correspondence to the parapet, i.e., WG22, as shown in Table 6 and Table 7. Although the value of ε is the dominant geometric variable, q

_{rear}also decreases with increasing h

_{n}(Figure 10). The best reduction was achieved for λ = h

_{n}/h

_{wall}≥ 0.3. Based on these literature results, 2 parapet configurations have been considered, i.e., ε = 30° and 45° and a fixed value of λ = 0.3. The resulting thickness of the nose w

_{n}is equal to 0.027 m and 0.046 m for ε = 30° and 45°, respectively.

_{rear}for the OBREC sections with and without the parapet. In the case of the straight crown wall, the greater the reservoir width the lower the overtopping at the rear side of the structure was. The presence of the nose reduces the values of q

_{rear}by 34% and by 41%, on average, for ε = 30° and 45°, respectively. For both inclinations, the configuration M3 shows the maximum reduction, i.e., 70% and 80% for ε = 30° and ε = 45°, respectively. Therefore, the inclusion of the nose leads to a safe harbor area even in case of a constrained reservoir width.

_{250}inside the reservoir and along the crown wall to understand the effects, in terms of pressures, of the presence of the nose. Due to the absence of overtopping at the rear side of the structure, the configuration M4 was not analyzed. The results of non-dimensional pressures acting on the crown wall and inside the reservoir are graphically represented in Figure 11 and Figure 12, respectively. In order to provide general results, the relative positions of the pressure transducer have been considered non-dimensionalised. In particular, in Figure 11, the vertical distances from the bottom (z) rather than the wave height (H

_{m0}) are shown. In Figure 12, the relative abscissa (x) of pressure transducers from the seaward border of the reservoir has been non-dimensionalised with respect to the reservoir width (B

_{r}).

_{r}.

_{rear}implied the parapet be inclined at 45° without increasing the pressures. The other geometric parameters, i.e., the berm and the reservoir width, do not affect q

_{rear}and p

_{250}.

## 4. Discussion of the Results with Reference to the Prototype Installation

_{reservoir}could increase the values of q

_{rear}. In particular, assuming that the water jets follow the tangent of the ramp crest, which works as a deflector, the greater the hydraulic performances of the frontal sloping ramp, the greater the value of up-rushing water driven directly on the upper part of the crown wall is. This means a potential higher overtopping discharge at the rear of the structure and a higher pressure on the parapet.

_{r}, and to meet the same spatial constraints, they have different dimensions of the reservoir width B

_{r}. The two values of R

_{r}have been selected according to the analysis of the typical wave climate of the site and to take into account the different turbine technologies, working with a different nominal head. The intention is not only to provide results for the sea climate at the study site but also to gather useful data easily exportable for more energetic sites. According to the analysis of the typical wave climate, the selected values of R

_{r}for NW-LAB and RS-LAB are 1.2 m and 2.0 m respectively (values referred to the mean tide level). The corresponding values of the reservoir width are 2.6 m and 3.7 m. The lower value of R

_{r}was selected according to the mean run-up of the most frequent wave (i.e., H

_{s}= 0.8 m). The upper value of R

_{r}was intended to capture the higher power generated by the higher waves. Its definition, however, followed a different approach. If one look at the maximum value of the power was multiplied by the wave frequency of occurrence (F), the prevalent wave height is associated to H

_{m0}= 2.2 m, T

_{p}= 6.8 s. In this case, however, the yearly average frequency of occurrence is too low (F = 1.4%) corresponding to just five days per year. Therefore, considering that the lowest nominal hydraulic head of the greatest low head Kaplan turbine [53] is 1.5 m, a ramp crest of about 2 m (1.78÷2.28 m for low- and high-water levels respectively) has been selected. This value can guarantee an average of about 30 equivalent working days per year of the RS-LAB (the name Real Scale, hence, indicates the site-specific considerations made for that ramp crest).

_{r}. The wider reservoir is more appropriate for the lower sloping plate, which is overtopped more frequently. The sensitivity of this parameter with respect to the hydraulic and structural performance, as it has been demonstrated during the experimental and numerical analysis, can be neglected. Therefore, the selection of B

_{r}can be done after the design of R

_{r}.

## 5. Conclusions

- to extend the experimental database of the laboratory tests performed in 2012 and in 2014 and
- to perform a sensitivity analysis related to the elements composing its cross section.

_{reservoir}or the pressure distribution along the structure. However, it slightly increases K

_{r}as expected, suggesting the design of a toe protection.

_{r}and p

_{250}, except for the wider reservoir that leads to an unstressed rear wall. Therefore, B

_{r}can be selected after the design of the sloping plate, which appeared to be the most sensitive element.

_{rear}by 34%, while the one inclined at 45° was reduced by 41%, on average. The maximum reduction has been measured for the configuration characterized by the smallest reservoir width, i.e., 70% and 80% for ε = 30° and ε = 45°, respectively. Therefore, the inclusion of the inclined parapet leads to a safer harbor area even in case of constrained values of B

_{r}. The best results have been obtained for the parapet inclined of 45°, which optimizes the values of q

_{rear}without increasing the pressures.

## Author Contributions

## Funding

## Acknowledgments

## Conflicts of Interest

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**Figure 1.**A 3-D rendering of a traditional rubble-mound breakwater and the Overtopping Breakwater for the Energy Conversion (OBREC).

**Figure 2.**(

**a**) The traditional rubble mound breakwater, (

**b**) the OBREC physical model, and (

**c**) the OBREC numerical model (adapted from Vicinanza 2014 [26], with permission from ELSEVIER, 2018). The main geometric parameters are: B

_{r}the reservoir width, B

_{s}the horizontal extension of the sloping plate, α its inclination, d

_{w}the height of the sloping plate and R

_{r}and R

_{c}the freeboard crests of the sloping plate and of the rear wall, respectively.

**Figure 4.**The positions of the water gauges to evaluate (1) the pressure acting on the device and (2) the q

_{reservoir}and q

_{rear}.

**Figure 6.**The numerical and theoretical vs. laboratory q*

_{reservoir}for the normal wave conditions.

**Figure 7.**(

**a**) The original d

_{w,high}configuration (M1); (

**b**) the configuration without a berm (M2); (

**c**) B

_{r}= 0.3 m (M3); (

**d**) B

_{r}= 0.5 m (M4); and (

**e**) the sloping plate 90° + 30° (M5).

**Figure 8.**A comparison of the values of K

_{r}for the M1–M5 configurations, under normal conditions only (Table 1).

**Figure 9.**The non-dimensional overtopping flow rate at the rear of the structure, q*

_{reservoir}, for the M1–M5 configurations compared with case M2 under normal conditions (Table 1).

**Figure 11.**The non-dimensional pressures on the crown wall for the configurations (

**a**) M1, (

**b**) M2, (

**c**) M3 and (

**d**) M5. The vertical distances from the bottom (z) rather than the wave height (H

_{m0}) are reported on the ordinate, while the statistical values of p

_{250}divided by the wave height (H

_{m0}), the density (ρ) and the gravitational acceleration (g), on the abscissa.

**Figure 12.**The non-dimensional pressures inside the reservoir for the configurations (

**a**) M1, (

**b**) M2, (

**c**) M3 and (

**d**) M5. The statistical values of p

_{250}divided by the wave height (H

_{m0}), the water density (ρ) and the gravitational acceleration (g) are reported on the ordinate; while the relative abscissa (x) of pressure transducers from the seaward border of the reservoir has been non-dimensionalised with respect to the reservoir width (Br), on the abscissa.

**Figure 14.**The cross sections of the OBREC prototype installed in the Naples harbor: (

**a**) the RS-LAB (Real Scale Laboratory) configuration and (

**b**) the NW-LAB (Natural Waves Laboratory) configuration.

Tests | h (m) | H_{m0} (m) | T_{m−1,0} (s) | R_{c} (m) | R_{r} (m) |
---|---|---|---|---|---|

Normal | 0.27 | 0.077–0.149 | 1.327–2.090 | 0.155 | 0.27 |

Extreme | 0.34 | 0.193 | 2.233 | 0.085 | 0.20 |

**Table 2.**The experimental vs. numerical q

_{reservoir}obtained by varying the porosities n assigned to the layers of the breakwater, i.e., armour, filter and core.

Configuration | Armour | Filter | Core | q_{res,exp} (L/m/s) | q_{res,num} (L/m/s) |
---|---|---|---|---|---|

1 | 0.8 | 0.7 | 0.6 | 0.046 | 0.073 |

2 | 0.7 | 0.6 | 0.5 | 0.046 | 0.056 |

3 | 0.6 | 0.5 | 0.4 | 0.046 | 0.006 |

4 | 0.7 | 0.5 | 0.4 | 0.046 | 0.004 |

**Table 3.**The geometrical characteristics of the configurations analyzed (Figure 7).

Configuration | Berm | R_{c} (m) | R_{r} (m) | B_{s} (m) | B_{r} (m) | α_{off, plate} (°) |
---|---|---|---|---|---|---|

M1 | √ | 0.27 | 0.155 | 0.3074 | 0.41 | 34° |

M2 | × | 0.27 | 0.155 | 0.3074 | 0.41 | 34° |

M3 | × | 0.27 | 0.155 | 0.3074 | 0.30 | 34° |

M4 | × | 0.27 | 0.155 | 0.3074 | 0.50 | 34° |

M5 | × | 0.27 | 0.155 | 0.3074 | 0.41 | 90° + 30° |

**Table 4.**The p

_{250}values acting on the sloping plate, the crown wall and the bottom part of the reservoir (uplift pressures) in kPa. See Figure 5 for the gauge locations.

WG No. | Lab | M1 | M2 | M3 | M4 | M5 |
---|---|---|---|---|---|---|

27 | 1.66 | 1.68 | 1.69 | 1.64 | 1.77 | 1.66 |

28 | 1.54 | 1.50 | 1.53 | 1.46 | 1.59 | 1.47 |

29 | 1.44 | 1.30 | 1.35 | 1.26 | 1.37 | 1.27 |

21 | 1.45 | 1.09 | 1.12 | 1.06 | 1.15 | 0.96 |

26 | 1.82 | 1.01 | 1.07 | 1.03 | 1.09 | 0.79 |

30 | 1.96 | 0.58 | 0.62 | 0.76 | 0.57 | 0.62 |

7 | 2.75 | 2.18 | 2.11 | 2.14 | 0 | 2.06 |

9 | 2.67 | 1.77 | 1.72 | 1.77 | 0 | 1.65 |

17 | 2.70 | 1.52 | 1.32 | 1.36 | 0 | 1.26 |

22 | 1.67 | 0.97 | 0.91 | 0.80 | 0 | 0.92 |

13 | 2.09 | 2.30 | 2.28 | 2.20 | 2.30 | 2.27 |

12 | 1.89 | 2.13 | 2.11 | 2.01 | 2.15 | 2.10 |

11 | 1.84 | 1.95 | 1.94 | 1.81 | 2.00 | 1.93 |

10 | 1.52 | 1.75 | 1.74 | / | 1.83 | 1.74 |

**Table 5.**The pressures acting inside the reservoir in kPa, with the same abscissa of the pressure transducers related to the uplift pressures.

WG No. | Lab | M1 | M2 | M3 | M4 | M5 |
---|---|---|---|---|---|---|

13_{in} | / | 1.71 | 1.73 | 1.84 | 1.75 | 1.77 |

12_{in} | / | 1.60 | 1.60 | 1.75 | 1.54 | 1.62 |

11_{in} | / | 1.80 | 1.83 | 2.06 | 1.64 | 1.82 |

10_{in} | / | 2.06 | 2.07 | / | 1.83 | 2.01 |

**Table 6.**The average overtopping discharge at the rear side of the crown wall (q

_{rear}(L/s/m)) for cases without parapet and with parapet (ε = 30° and 45° respectively).

M1 | M2 | M3 | M4 | M5 | |
---|---|---|---|---|---|

Crown wall | 0.47 | 0.45 | 0.63 | 0.00 | 0.44 |

Crown wall and parapet (ε = 30°) | 0.28 | 0.30 | 0.19 | 0.00 | 0.31 |

Crown wall and parapet (ε = 45°) | 0.26 | 0.25 | 0.13 | 0.00 | 0.28 |

**Table 7.**Thee p

_{250}values inside the reservoir and on the crown wall (kPa), for ε = 0° (no parapet), ε = 30° and ε = 45°. The positions of the water gauges (WG) are shown in Figure 4.

Wg | M1 | M2 | M3 | M5 |
---|---|---|---|---|

ε = 0° | ||||

7 | 2.18 | 2.11 | 2.14 | 2.06 |

9 | 1.77 | 1.72 | 1.77 | 1.65 |

17 | 1.52 | 1.32 | 1.36 | 1.26 |

22 | 0.97 | 0.91 | 0.8 | 0.92 |

13in | 1.71 | 1.73 | 1.84 | 1.77 |

12in | 1.6 | 1.6 | 1.75 | 1.62 |

11in | 1.8 | 1.83 | 2.06 | 1.82 |

10in | 2.06 | 2.07 | / | 2.01 |

ε = 30° | ||||

7 | 2.13 | 2.14 | 2.32 | 2.17 |

9 | 1.76 | 1.75 | 1.93 | 1.76 |

17 | 1.43 | 1.4 | 1.56 | 1.41 |

22 | 1.49 | 1.58 | 1.28 | 1.52 |

13in | 1.72 | 1.73 | 1.95 | 1.73 |

12in | 1.59 | 1.6 | 1.91 | 1.62 |

11in | 1.8 | 1.8 | 2.22 | 1.83 |

10in | 2.07 | 2.08 | / | 2.08 |

ε = 45° | ||||

7 | 2.39 | 2.18 | 2.4 | 2.23 |

9 | 2.05 | 1.8 | 2.01 | 1.85 |

17 | 1.88 | 1.44 | 1.64 | 1.57 |

22 | 2.38 | 1.54 | 1.57 | 1.83 |

13in | 1.76 | 1.74 | 2 | 1.76 |

12in | 1.69 | 1.62 | 1.92 | 1.67 |

11in | 1.91 | 1.84 | 2.27 | 1.89 |

10in | 2.26 | 2.12 | / | 2.15 |

© 2019 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

## Share and Cite

**MDPI and ACS Style**

Palma, G.; Mizar Formentin, S.; Zanuttigh, B.; Contestabile, P.; Vicinanza, D.
Numerical Simulations of the Hydraulic Performance of a Breakwater-Integrated Overtopping Wave Energy Converter. *J. Mar. Sci. Eng.* **2019**, *7*, 38.
https://doi.org/10.3390/jmse7020038

**AMA Style**

Palma G, Mizar Formentin S, Zanuttigh B, Contestabile P, Vicinanza D.
Numerical Simulations of the Hydraulic Performance of a Breakwater-Integrated Overtopping Wave Energy Converter. *Journal of Marine Science and Engineering*. 2019; 7(2):38.
https://doi.org/10.3390/jmse7020038

**Chicago/Turabian Style**

Palma, Giuseppina, Sara Mizar Formentin, Barbara Zanuttigh, Pasquale Contestabile, and Diego Vicinanza.
2019. "Numerical Simulations of the Hydraulic Performance of a Breakwater-Integrated Overtopping Wave Energy Converter" *Journal of Marine Science and Engineering* 7, no. 2: 38.
https://doi.org/10.3390/jmse7020038