Monopile Fixed Foundation

This section assesses the methodology of design of a monopile foundation and provides results from a parameterised model developed using Morison’s Equation for hydrodynamic loads from spectral waves as well as turbulent wind loads at the hub for the three turbine types described above. The method of superposition, as first introduced in hydrodynamics by [114] is applied to both the wave spectra and wind spectra to provide a time series of representative loading on the structure. The variation of structural mass of the monopile foundation will be investigated for increasing water depth, turbine size and significant wave height and peak wave period. Only loading from spectral waves on the foundation and stochastic wind on the turbine rotor are considered for design of the monopile and TP using bending moment and shear checks in accordance with [116] as referred from [115]. A number of assumptions are made in the model and are outlined below.

• The thrust curve for the NREL 5MW reference turbine has been divided by the turbine rating to provide an approximation of thrust on the turbine hub depending on the rating of the turbine. This approximation may be updated with a correct thrust curve for the turbine under study.

• Only mean wind speeds of 12m/s are considered in the results below to allow for maximum thrust at the hub.

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• The water depth determines the maximum significant wave height possible through

𝐻𝑏 = 0.42𝑑𝑏 Equation 3-1

where Hb is the breaking wave height and db is the breaker depth. The breaker depth is assumed to be the water depth under study. The maximum Hs is limited to 15m.

• The design wave period is determined from the design wave height and is the minimum period allowed before breaking, thus giving the steepest wave possible to achieve maximum loading.

𝑇𝐷 = 11.1�𝐻_𝑏

𝑔 Equation 3-2

• The hub height of the turbine is assumed to be constant at 90m.

• The material factor for steel is constant at 1.15 as stated in [115].

• The yield strength of steel is 235MPa, representative of NS from [115].

• Monopile diameter is constant at 6m.

• The point of fixity of the monopile is assumed to be 5m below the seabed level.

• The drag and inertia coefficient for Morison’s Equation are assumed to be 0.65 and 1.6 respectively as outlined in [117]

• No current profile is used in the calculation of hydrodynamic loads.

Therefore the study is primarily driven by the choice of, 1. Water depth,

2. Turbine rating,

3. Monopile steel plate thickness.

3.1.1.1 Results

This section presents the results of the study on the monopile foundation, beginning initially with a sample of the outputs from the parameterised model.

3.1.1.1.1 Sample Output Plots from Model

In terms of the geometry of the monopile foundation, it was outlined that the diameter of the pile was kept constant. Two further parameters of the monopile geometry were required, the height above sea level and the depth of penetration below the seabed. These two parameters have been taken from a sample provided in [117]. The sample has been used to formulate the platform level and pile penetration depth as a function of water depth in which the foundation is installed. These are illustrated in Figure 3-1and Figure 3-2.

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Figure 3-1: Platform Level as a Function of Water Depth

Figure 3-2: Monopile Penetration Depth as a Function of Water Depth

Following the determination of the pile geometry, the calculation of the loading was determined. As the model was required to analyse loading from spectral waves and stochastic wind, both a wave and wind spectrum was required as an input. The wave spectrum used was the Bretschneider Spectrum and the wind spectrum used was the Kaimal Spectrum as recommended by [118], while the Reference Turbulence Intensity Factor is given as a function of height above sea level by [119]. The

“principle of superposition” has been applied to both spectra to create a representative time series of wave conditions based on linear wave theory and wind speeds. This principle involves determining the amplitude of the sinusoidal signal for

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each frequency within the relevant spectrum and attributing a random phase shift to each sinusoid. The water level elevation, velocity and acceleration of a sample wave condition are illustrated in Figure 3-3. This process is carried out at depth increments of 1m so as a time series of velocity and acceleration can be determined at each depth. This allows the variation in loading with water depth to be applied to the foundation. The rationale behind the use of basic unmodified linear theory is to provide a simplified estimate of the order of magnitude of the loads which the structures are likely to experience. As the design progresses for any structure, more refined and detailed analysis techniques, theories and models should be used. The extreme wave conditions are potentially better represented using non-linear wave theories, but these require much more effort and time to implement. As a first approximation, linear theory is deemed reasonable to apply.

The instantaneous load on the foundation at each depth as determined from Morison’s Equation is applied and the summation of these gives the lateral force applied to the foundation by the incoming waves. A sample output of the lateral force from wave loading applied at the point of fixity of the pile is illustrated in Figure 3-4. The associated bending moment at the point of fixity of the pile is then simply the force times the distance from the point of fixity. A sample output of the bending moment from wave loading is illustrated in Figure 3-5.

Figure 3-3: Sample Time Series of Water Level Elevation, Velocity and Acceleration at 5m Below Surface

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Figure 3-4: Sample Instantaneous Lateral Force at the Point of Fixity of the Monopile from Wave Loading

Figure 3-5: Sample Instantaneous Bending Moment at the Point of Fixity of the Monopile from Wave Loading

The other source of loading considered in this study is that of the thrust applied to the wind turbine at the hub. The hub height is constant at 90m above sea level while the thrust is determined from the wind speed, thrust per MW capacity curve and wind turbine rating. The instantaneous wind speed is determined from the Kaimal

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Spectrum and superposition as discussed. Figure 3-6 illustrates a sample wind speed time series at hub height used in the model.

Figure 3-6: Sample Instantaneous Wind Speed at Hub Height of the Turbine

The wind speed time series is then used in conjunction with the thrust curve to determine the instantaneous lateral thrust load applied to the turbine at the hub height. A sample thrust time series is illustrated in Figure 3-7. It may be noticed that the thrust reaches a maximum value irrespective of the wind speed. This is due to the fact that the thrust curve is formed from an average thrust value for an averaged wind speed. In reality, the thrust at the rated wind speed may be higher depending on the speed at which the control mechanisms of blade pitch angle can react. For the purposes of this preliminary study, this method will suffice.

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Figure 3-7: Sample Instantaneous Wind Thrust at Hub Height of the Turbine

Figure 3-8: Sample Instantaneous Bending Moment at the Point of Fixity of the Monopile due to Wind Loading

The bending moment applied to the point of fixity of the pile may then be calculated by multiplying by the distance between the point of fixity of the pile and the hub height as illustrated in Figure 3-8. The contribution of both wind and wave loading on the pile are summed and the total shear force and bending moment applied to the pile at the point of fixity is determined. A sample plot of these is illustrated in Figure 3-9.

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Figure 3-9: Sample Total Lateral Force and Bending Moment on the Monopile

The steel plate thickness from which the monopile must be fabricated to resist this loading is designed according to [116] following suitable load factors being applied to the design shear force and moment.

3.1.1.1.2 Model Validation

Simple tank testing has been carried out on a number of hybrid concepts to be discussed further in Chapter 4. One such model was a long slender pile with load cells incorporated into the pile to measure the hydrodynamic loading from incident waves described in APPENDIX C. The mathematical model described above has been used to estimate the loads likely to be experienced by the pile. The model configuration, in prototype scale, was a pile diameter of 3m in a water depth of 30m.

The wave conditions were 3m H_s and 10s T_p Bretschneider Spectrum. The comparison between measured and modelled results is shown in Table 3-1.

Table 3-1: Validation of Mathematical Model against Measured Data

Model (kN) Measured (Scaled Up) (kN)

% Difference

Load @ -10m 8.3 9.5 -12.5

Load @ -20m 5.1 7.1 -28

The results indicate that the mathematical model based on Morison’s Equation underestimates the loads at both levels below water level. This may be due to several reasons including, incorrect drag and inertia coefficients used in the model, proportionately larger loads due to viscous effects at small scale tank testing which are then carried through in the scaled up figures. This is a common feature of tank testing in that such loads and damping are over-estimated relative to prototype due to scale effects of viscous forces and vortex shredding. As the figures are not an order

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of magnitude larger than one another, this model may be deemed sufficient for preliminary design purposes.

3.1.1.1.3 Vestas V90 3MW Turbine Monopile Sizes

Figure 3-10 illustrates the dependency of monopile steel mass on water depth for the Vestas 3.0MW turbine. The results can be assumed to follow a linear trendline which is included in the plot as well as the equation for interpolation of further sizes depending on water depth.

Figure 3-10: Dependency of Monopile Total Mass on Water Depth for 3.0MW Turbine

3.1.1.1.4 Siemens SWT 3.6-120 3.6MW Turbine Monopile Sizes

Figure 3-11 illustrates the dependency of monopile steel mass on water depth for the Siemens 3.6MW turbine. The results, similar to the 3.0MW turbine, may be assumed to follow a linear trendline which is plotted for illustrative purposes as well as the line equation for interpolation of other monopile sizes depending on the water depth of interest.

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Figure 3-11: Dependency of Monopile Total Mass on Water Depth for 3.6MW Turbine

3.1.1.1.5 NREL 5MW Turbine Monopile Sizes

Figure 3-12 illustrates the dependency of monopile steel mass on water depth for the NREL 5.0MW turbine. The result is slightly different from the previous turbines as the lateral forces from the turbine are higher and contribute more to the steel requirements of the pile in lower water depths. As the water depth increases, the pile steel masses become similar to that required for the 3.6MW turbine in Figure 3-11as wave loading again becomes the dominant force. The dependency of the pile mass for the 5.0MW turbine is a cubic function in this case and the equation is included in the plot for estimation of pile sizes in other water depths of interest.

3.1.1.2 Comments on Results

The results illustrate the trends to be expected in the design of monopile foundations for offshore wind farms, both existing and future farms. For the case of existing farms, take for example the results for the 3.6MW turbine in Figure 3-11. The foundation steel mass in shallow water of 15m is ~200t, while in 30m water depth this increases fourfold to ~800t while the depth has only doubled. The design appears to be weakly dependent on turbine size with the exception of designs in depths lower than 30m when the larger turbines provide significantly higher thrust loads and therefore bending moments at the foundation. It is anticipated that as the depth becomes higher that the foundation design will be dominated by wave loading and the turbine size will have a relatively weak effect. The results confirm the current train of thought surrounding monopile foundations by the industry. They simply cannot be feasible in deep water sites given the rate of increase of steel mass dependency on water depth. This is the reason for rigorous increase in efforts in design of alternatives including gravity based and steel jacket foundations.

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Figure 3-12: Dependency of Monopile Mass on Water Depth for 5.0MW Turbine

3.1.1.3 Unit Costs of Monopile Foundations

As the monopile foundation is the most widely used type, a generally accepted average unit cost per tonne production is approximately €2000/tonne steel mass.

This is however a function of the steel plate sizes and the length of the monopile, as these imply the number of welds likely to be carried out during manufacturing which dictates the manufacturing costs. This section provides a means to estimate the variation in unit costs of monpiles using confidential information for a selected case, namely the 3.6MW turbine in 20m water depth. The case presumes the following details in Table 3-2.

Table 3-2: Base Case Monopile Foundation Details

Item Value

Monopile Mass 254.4 t

TP Mass 118 t

Pile Diameter 6 m Steel Plate Thickness 35 mm Monopile Length 49 m

TP Length 22.9 m

From a confidential percentage breakdown of the cost of manufacturing a monopile foundation, it is found that the approximate split of unit costs of a monopile and TP is 91% and 117% of the average respectively based on the Belwind Offshore Wind Farm average monopile and TP mass ratio of 30%. The market price of steel plate

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and hot rolled sections is approximately €600/t currently [120]. Thus the fabrication cost of a monopile is ~€1230/t. Back calculating using these figures for the case study of the 3.6MW turbine in 20m water depth, it is found that the labour and overhead cost per hour is ~€31.20/hr based on the breakdown of fabrication costs from the [121], outlined in Table 3-3.

Table 3-3: Steel Fabrication Costs Breakdown

Welding Costs Breakdown % Labour and Overheads 85

Materials 9

Equipment Investment 4 Power and/or Gas 1

This cost of labour is determined from the time taken to weld all “cans” of a monopile together using standard steel plate sizes. The welding rates and operation times are taken from [122]. The preparation of a steel “can” for a monopile is illustrated in Figure 3-13 while a finished pile is illustrated in Figure 3-14. The rate breakdowns are used in the estimation of the unit costs of the other monopile sizes as calculated from the previous study.

Figure 3-13: Monopile “Can” Rolling

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Figure 3-14: Completed Monopile

Figure 3-15 illustrates the dependency of the fabrication cost of the monopile on water depth for a 3.0MW turbine. Similar to the dependency of steel mass on depth, the unit cost of production has a linear dependency on depth.

Figure 3-15: Unit Costs of Monopile Only as a Function of Water Depth for 3.0MW Turbine

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Figure 3-16: Unit Costs of Monopile Only as a Function of Water Depth for 3.6MW Turbine

Figure 3-16 illustrated the unit cost of production of a monopile dependency on water depth for a 3.6MW turbine. Again, a linear dependency is observed for the unit cost with water depth similar to the steel mass dependency.

Figure 3-17: Unit Costs of Monopile Only as a Function of Water Depth for 5.0MW Turbine

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It may be seen that the unit cost of the monopile is only weakly dependent on the turbine size, and mainly dictated by the water depth to be deployed in. The nature of the dependency of both monopile and TP mass and the unit costs of same can be seen to become more non-linear as the turbine size gets larger as illustrated in Figure 3-17 for the 5MW WT. This is similar to the mass dependency on depth and is likely to become linear again in deeper water depths as wave loading dominates. For the purposes of this thesis, these are the limits which will be investigated however, in the future as larger turbines come to market, the dependency of the mass and unit costs of the monopile foundation should be investigated and alternatives considered. One final observation, from both the structural mass dependency and unit cost of production dependency on depth is that as the depth gets larger, the piles become larger, heavier and progressively more expensive to produce.