AVATAR Reference Blade Design

AVATAR Deliverable D1.2 Reference Blade Design

January 30, 2015 Agreement n.: FP7-ENERGY-2013-1/ n° 608396 Duration: November 2013 to November 2017 Coordinator: ECN Wind Energy, Petten, The Netherlands

Supported by:

Page 2 of 52 WP1.2: Design for the AVATAR RWT Rotor

Document information

Document Name: AVATAR Reference Blade Design

Confidentiality Class XX

Document Number: D1.2

Editor: G. Sieros (CRES)

Contributing authors: D. Lekou, D. Chortis, P. Chaviaropoulos (CRES)

X. Munduate, A. Irisarri (CENER)

H. Aa. Madsen, K. Yde, K. Thomsen(DTU)

M. Stettner, M. Reijerkerk (GE)

F. Grasso, R. Savenije, G. Schepers (ECN)

C.F. Andersen (LM)

Review: G. Schepers (ECN)

Date: 30/01/2015

WP: WP1 - Integration and Evaluation of 10MW Rotor

Task: Task1.2 Design for the AVATAR RWT Rotor


Table of contents

1 Introduction 4

2 Specifications 5

3 Airfoil Selection 7

4 Planform Design 11

5 Structural Design 16

5.1 Initial Design 16 5.1.1 Geometry 16 5.1.2 Loads 16 5.1.3 Static Analysis 18 5.1.4 Modal Analysis 23 5.1.5 Buckling Analysis 23 5.1.6 Mass & Stiffness Distribution 24 5.2 Redesign of the blade 25 5.2.1 Initial Design for weight estimation 27

6 Loads and Stability 40

6.1 Test Cases 40 6.2 Computational Tools 40 6.3 Results 42 6.3.1 Power Curve Verification 42 6.3.2 Speed-up/Speed-down simulations 45 6.3.3 Unsteady operation 47

7 Conclusions 51

8 References 52


1 Introduction This document represents Deliverable D1.2 of the AVATAR project, containing the design

process followed for the reference wind turbine. The design is based on the INNWIND.EU

10MW research wind turbine, with modifications in order to explore more demanding flow

regimes. A detailed description of the resulting geometry and structural properties is also part of

the deliverable. Throughout this document, the “baseline 10MW wind turbine” referred to is the

INNWIND.EU research wind turbine / DTU 10MW RWT (Bak, et al.)


2 Specifications Even though it was not possible (or required) to have a production-quality design for the project,

a connection to standard industry practices was maintained. The industrial partners provided the

general guidelines for the blade design that would need to be satisfied. These included

manufacturability constraints and a minimum set of IEC load cases to be checked at the initial

design stage. Details of the constrains that were used are given in (Stettner & Chaviaropoulos,

2014)

It is useful to introduce the background for the design that is proposed. We use the standard

definitions for Tip-Speed-Ratio (TSR or λ), blade section lift to drag ratio k and the radius

fraction x according to Eq. (1). We denote V the ambient wind speed, ω the rotational speed

and R the rotor radius and B is the number of blades.

𝜆 = 𝜔𝑅/𝑉 𝑘 = 𝐶𝐿/𝐶𝐷 𝑥 = 𝑟/𝑅 (1)

We introduce the non-dimensional lift distribution Λ(λ,x) as

𝛬(𝜆, 𝑥) =𝑐(𝜆, 𝑥)𝐶𝐿

𝑅 (2)

where c(λ,x) is the chord distribution and CL the lift coefficient.

For a pitch-controlled, variable speed HAWT design and for a given rotor radius the classical

rotor aerodynamic design problem would seek to maximize the energy capture by maximizing

the power coefficient CP. According to the BEM theory this would happen for an axial induction

value α=1/3 and would correspond to a TSR design value λ which gets larger (along with CP,MAX)

as the aerodynamic performance of the blades k gets better (higher). As design λ increases the

non-dimensional lift distribution Λ(λ,x) gets smaller and, for the same family of blade profiles, the

rotor solidity gets lower.

For a variable speed rotor, the design λ value (and therefore CP,MAX) can be maintained over a

range starting from a minimum wind speed, defined by the low-end capability of the variable-

speed power conversion system, up to a maximum wind speed which is limited by the rotor

maximum tip-speed, either for restraining noise or centrifugal loading. We shall call this

maximum wind speed, where CP = CP,MAX , “design wind speed”. Usually, the pitch variable

speed turbines have their design wind speed just below their rated speed. The left of Figure 1

presents a plot of CP versus α for a three-bladed rotor with profiles of k=100. The CP,MAX value is

0.4966 corresponding to an α value of 1/3.

Suppose that we have an initial (reference) wind turbine and we want to add some freedom in

our design by redesigning the rotor, letting its radius free, but respecting all turbine related

constrains (the rated rotational speed and power, the hub loading etc). We will assume for

simplicity that the new rotor will use the same family of airfoils (same k, considering Reynolds

number effects as secondary at the scale of our interest).


Let R0 be the initial rotor radius and let subscript “0” denote the reference design, the one with

α=1/3 corresponding to maximum CP.

The new design problem is formulated as:

𝐶𝑃(𝜆,𝛼)⋅𝑅2

𝐶𝑃0(𝜆0,𝛼0)⋅𝑅02 → 𝑚𝑎𝑥,

𝑠𝑢𝑏𝑗𝑒𝑐𝑡 𝑡𝑜 𝐶𝑀(0)(𝜆, 𝛼) ⋅ 𝑅3

𝐶𝑀0(0)(𝜆0, 𝛼0) ⋅ 𝑅03 ≅ 1

(3)

That is: “maximize the power output up to the design wind speed without exceeding the initial

aerodynamic root blade moment”. By eliminating the radius dependence the optimization

problem (3) can be recast as:

𝐶𝑃(𝜆, 𝛼)

𝐶𝑀(0)(𝜆, 𝛼)2

3⁄→ 𝑚𝑎𝑥 (4)

The solution of the optimization problem (4) for α, given λ=-λ0 , is shown at the right of Figure 1.

The resulting value is α=0.187 and although it has been calculated for a “wrong” λ (= λ0) it can

be shown that the optimum α solution for (4) is more or less λ insensitive.

As a result, the optimal rotor will have a larger radius; will capture more energy at its design

conditions; and will be less loaded than the initial one (design CT and CM(r) will be smaller),

operating at a lower axial induction value α~0.20 instead of α0=0.33. In other words, we can

sacrifice CP in order to increase energy capture with a larger rotor diameter, while maintaining

the aerodynamic bending moments at their initial level. This is feasible thanks to the special

shape of the CP and CM(0) curves, where moving a little left from the optimum α at the CP curve

the power coefficient loss is milder than the corresponding reduction of bending moments in the

CM plot.

This analysis summarizes the justification for moving to higher tip speeds and larger diameters

for the proposed design.

Figure 1 Plots of non-dimensional coefficients, candidates for blade optimization, versus axial induction coefficient α


3 Airfoil Selection The geometrical characteristics of the airfoils were not fixed at the specifications level, as they

were going to be part of the design procedure. It is expected that because of the unusual

aerodynamic characteristics of the wind turbine, existing airfoil families are probably suboptimal

for the final design. A new set of airfoil families will be produced later in the course of the

project, but for the reference wind turbine, an existing family needs to be utilized. The main

requirements were that it is readily available, and that some experimental data for operation at

various Reynolds numbers are also available (though further tests at high Re numbers will be

performed in the context of the project).

It was therefore decided to use the established

DU airfoil families at thicknesses shown in

Error! Not a valid bookmark self-reference..

The main challenge lies in acquiring reliable

polar curves for theses airfoils at the very high

Re numbers involved. As the planform design

was not finalized at the time of the airfoil

selection, an estimate was performed

assuming chords similar to the ones used for

the INNWIND.EU rotor. The proposed sections

are based on the following assumptions

1. We assume that (as a first approximation)

the INNWIND.EU rotor is scaled in the

radial direction, but not in the other

directions, so that the chord lengths remain

similar (this is compatible with the

requirement to retain bending moments similar to the INNWIND rotor)

2. A scaling factor of 1.15 is applied in the radial direction

3. The rotating speed is the same as for the INNWIND.EU rotor (9.6RPM at rated, 6RPM

minimum)

4. A similar hub → tip thickness distribution is used

5. The twist distribution is not known at this moment, but is not important for the polars

calculation

Based on these simplifying assumptions, a preliminary distribution of the sections would be as

shown in the following table. This is not necessarily an accurate description of the final

geometry, but as we are only interested in defining the approximate Reynolds number, it should

be sufficient.

Table 1 Airfoil sections used for the reference blade

60.0% Artificial based on thickest available DU

40.1% DU 00-W2-401

35.0% DU 00-W2-350

30.0% DU 97-W-300

24.0% DU 91-W2-250 (modified for t/c=24%)

21.0% Based on DU 00-W-212 – added trailing edge thickness


Table 2 Estimate of airfoil operating conditions

Radius [%] 17 25 32 40 55 95 Chord [m] 5.7 6.1 6.2 5.8 5 1.8 t/c [%] 60 40 35 30 21-24 21-24

The resulting design conditions are summarized in Table 3. The 24% and 21% airfoils will be

used for a larger part of the blade, so conditions characteristic for 50%, 75% and 95% of the

radius are included.

Initial polars were estimated using the RFOIL computational method developed at ECN. Full

360deg polars were produced for all the design conditions, so that they can be used for

aeroelastic as well as performance calculations.

Table 3 Design conditions for the airfoils used for the AVATAR RWT

Section Thickness

Re (rated) Ma (rated) Re (Min) Ma (Min)

60.0% 7.0×106 0.05 4.4×106 0.03

40.1% 11.0×106 0.07 7.0×106 0.05

35.0% 14.0×106 0.09 9.0×106 0.06

30.0% 17.0×106 0.12 10.0×106 0.07

24.0% 20.0×106 0.16 12.0×106 0.10

24.0% 16.0×106 0.25 11.0×106 0.15

24.0% 13.0×106 0.30 8.0×106 0.18

21.0% 20.0×106 0.16 12.0×106 0.10

21.0% 16.0×106 0.25 11.0×106 0.15

21.0% 13.0×106 0.30 8.0×106 0.18

The results for the different airfoil families are given in Figure 2 - Figure 7. The performance

levels have been adjusted to be consistent with expected values at lower Reynolds numbers (to

compensate for a tendency to overpredict the performance). A trailing edge thickness from 0.6-

1.5% is used, in-line with the values used in INNWIND.EU and the specifications for

manufacturability.


Figure 2 Lift and drag characteristics for t/c=21% (three Re numbers)

Figure 3 Lift and drag characteristics for t/c=24% (three Re numbers)

Figure 4 Lift and drag characteristics for t/c=30%






4 Planform Design For the given airfoil families, an optimized spanwise distribution of the chord, thickness and twist

Figure 9) were obtained through an iterative process. The reference blade from the

INNWIND.EU project (Table 4) was used as a starting point, with modifications respecting the

general specifications agreed upon for the AVATAR blade (Stettner & Chaviaropoulos, 2014).

The constraints on the resulting blade were:

The specific power should be drastically

reduced (from 400W/m2 → 300W/m2). This goes

together with an increase of blade radius from

89 to 102.5 meter.

Thrust decreased, so that bending moment at

tower bottom remains constant

Rotational speed should be kept the same (i.e.

a rated value of 9.6 rpm), so that the other

subcomponents will be interchangeable

between the two wind turbines.

The result is a low induction blade, where different

operating strategies can be employed at part load

operation, either “peak-shaving” the root bending

moment through pitching, or by keeping a constant low

induction at partial load (see Figure 8).

For the derivation of the optimum planform design a

constrained optimization problem is setup, where the free variables are:

Chord length at 3-4 different positions along the span. Based on the specifications, there

are limits on maximum chord and chord at the hub (in order to accommodate the

pitching mechanism).

Blade twist value at 3-4 different positions along the span.

Blade thickness and position where thickness switches to minimum value

Design Tip Speed Ratio

The design constraints that were based on the previous discussion about retaining the load

levels remain unaffected.

Keep the blade root bending moment the same as for the INNWIND.EU blade. This

results in the lowest loading, with very low induction factors along the blade

Keep the blade thrust the same as for the INNWIND.EU blade. Using this solution the

blade root bending moment is allowed to increase somewhat, leading to designs with

induction factors in the range 0.20-0.25

Keep the tower bottom bending moment same as for the INNWIND.EU blade.

Figure 8. Blade root bending

moment when using constant low induction (red) or peak shaving (blue)


Table 4 INNWIND.EU wind turbine general characteristics (Bak, et al.)

Wind Regime IEC Class 1A

Rotor Orientation Clockwise rotation - Upwind

Control Variable Speed

Collective Pitch

Cut in wind speed 4

Cut out wind speed 25

Rated wind speed 11.4

Rated power 10

Number of blades 3

Rotor Diameter 178.3

Hub Diameter 5.6

Hub Height 119.0

Minimum Rotor Speed 6.0

Maximum Rotor Speed 9.6

Maximum Generator Speed 480.0

Gearbox Ratio 50.0

Maximum Tip Speed 90.0

Different optimization strategies were tried at ECN and CRES. In the simplest case the pitch

angle was constant up to rated speed and only allowed to vary from there on, in order to keep

the power output constant. In some of the examined cases, the pitching schedule was allowed

to vary in the benefit of power capture but, also, with increased mean flapping loads

Figure 9 Planform design for the reference wind turbine


A comparison between the initial INNWIND.EU blade, the proposed AVATAR blade and the

SANDIA 100m blade (Griffith & Ashwill, The Sandia 100-meter all-glass baseline wind turbine

blade: SNL100-00, 2011) (Griffith, The SNL-100-01 Blade: Carbod Design Studies for the

Sandia 100m Blade, 2013), which is designed for a 13MW generator. As seen in Figure 12 the

chord distribution is similar between the two designs, while a substantially larger chord is used

for the Sandia blade. The actual (absolute) thicknesses of the blades are quite close to each

other.

The linear density of the Sandia blade is also larger, resulting in a substantially heavier blade

(expected because of the 30% larger power). Nevertheless, and in spite of the constraint for

constant blade loading, the AVATAR blade unavoidably has a higher fatigue loading, compared

to the INNWIND.EU blade (Figure 13).

Figure 10 Spanwise distribution of properties for different wind speeds. The results for 8m/s and 10m/s are before the start of the pitching action, showing approximately constant coefficients and axial induction. The final result is after commencement of pitching action.

8m/s

10m/s 12m/s


Figure 11 Estimated power curve and pitching schedule

Figure 12 Chord (left) and absolute thickness (right) comparison between Sandia, INNWIND.EU and AVATAR blades


Figure 13 Linear density comparison (left) and resulting circular loading based on own weight (right).

Figure 14 Rendering of the proposed blade 3D geometry


5 Structural Design An iterative procedure, based on the AVATAR specifications document (D1.1) was used for the

blade structural design. The major requirements for which a compliant structural solution (given

the geometry of the AVATAR blade) was sought were related to

Acceptable natural frequencies for the blade

Keeping the mass of the blade as close as possible to the INNWIND.EU reference blade

Keeping the limit of the flapwise bending deflection as close as possible to the DTU

reference blade

Strength and stability criteria should be complied with for the AVATAR blade (under an

initial estimated ultimate loading condition).

5.1 Initial Design The structural design was performed through 3D- shell finite element model, modelling the

composite material and the internal structure of the blade.

The solution includes spar caps with large percentage of unidirectional carbon fibre layers. The

outer shell is made of glass epoxy (unidirectional and +/-45 layers) using sandwich (balsa core)

constructions in trailing and leading edge. Instead of minimum usage of carbon to replace

unidirectional glass, so as to achieve the required stiffness the blade was designed from the

start to reach compliant properties (mass, stiffness, strength, structural stability) with carbon on

the spar caps. To also take into account manufacturability it was assumed that the blade outer

shell surface would be manufactured independent from an internal box beam (i.e. the shear

webs and the caps).

Material properties were taken from the reference DTU blade (and the INNWIND.EU

benchmark) and the relevant carbon/epoxy UD properties used in the initial estimations for the

AVATAR blade. Material strength and elasticity properties used in the model include the

relevant partial safety factors as prescribed by the design standards (IEC 61400-1) and

wherever necessary the DNV and GL design guidelines.

5.1.1 Geometry

The external geometry of the blade (and the aerodynamic planform) was given based on the

planform design (see previous section). The external geometry was given at 50 sections along

the blade. The blade root is at 2.8m from the hub centre, while the outer radius is at 102.88m.

5.1.2 Loads

For the initial structural design of the blade an initial estimation of the loads was done using info

from the reference DTU blade. The bending moments were increased by 10% for the AVATAR

blade. Some load cases were provided for the reference DTU blade. Out of these the worst

case of the resultant bending moment in the flap direction was used, based on the procedure for

selecting the worst design case during blade testing for certification. The loads for the DTU

reference blade are provided on each section with respect to the (local) primary elasticity axis.

These were then turned into the global blade axis using the (local) information of each section


for the structural pitch (i.e. the angle of the primary elasticity axis with respect to the chord of the

section) and the twist angle.

Nevertheless, assessing the result of the analysis showed that there are sharp increases and

decreases of the bending moment distribution along the blade length, which were not expected,

especially in the inboard area (between the section of the maximum chord and the root).

Therefore the bending moment distribution was smoothed using different structural pitch angle

than those provided for the reference DTU blade.

The final result for the estimated bending moment distribution on the global blade axis are

shown in Figure 15 and Figure 16 for the DTU reference blade for the flap and edge (lead-lag)

direction respectively, denoted “DTU case”.

From these loads a distribution was selected for the AVATAR blade by increasing the bending

moments by 10%. These bending moments distributions are shown on the graphs (Figure 15

and Figure 16) for the flap and edge direction respectively denoted as “AVATAR increased”. For

the flap bending moment distribution this result was implemented in the finite element model

used for the initial structural design of the blade. The distribution in Figure 15 is denoted “used

for design”. For the edgewise bending moment the distribution was even further smoothed to

attain a better distribution of loads along the blade length. The smoothing was done on the

conservative side. The bending moment distribution in the edge direction employed in the finite

element model is denoted in Figure 16 as “used for design”.

Figure 15 Bending moment in the flap direction


Figure 16 Bending moment in the edge direction

5.1.3 Static Analysis

The finite element analysis of the AVATAR blade yielded the following results:

Mass: 46347.2Kg (compare with about 42400kg of the DTU blade)

Center of gravity: 33.598m

Flapwise deflection: Uy=8.738m (compare with about 18m of the DTU blade)

Edgewise deflection: Ux=3.829m (compare with about 3.2m of the DTU blade)

Failure is predicted on single elements with a Tsai-wu failure criterion value of less than

1.1 (of course above 1.0). Since the loads are only an initial estimation we have allowed

some localized failures.

A summary of local failure point is shown in the figures, where the elements in white color

undergo failure. In Figure 17 the failure is located at 3rd section (A) and 11th section (B). Near

the tip (Figure 18), the failure is attributed to the flapwise force and it is located at the 50th

section (last section). Regarding the shear Webs the failure is located at 10th section (C), Figure

19. Near the tip, the failure on the webs is located at 40th section (D), 45th section (E) and 50th

section (F). All these failures can be avoided by adding a layer of BIAX along the webs’ length.


Figure 17 Local failure points (Tsai-Wu)


x

y

z

A

B

x

y

z




x y

z

C

B

x y

z

E

B

F

B

D

B


Figure 21 Strain distribution (EXX)

Figure 22 Strain distribution (EYY) x

y

z

x

y

z

x

y

z


Figure 23 Strain distribution (EXY)

x

y

z

x

y

z

x

y

z


5.1.4 Modal Analysis

The results of the natural frequencies by the modal analysis performed (using the 3-D shell

element model) are given below. Note that the rotating frequency of the wind turbine is

1P=0.16Hz.

§ 1st Eig: 0.639Hz (Edge)

§ 2nd Eig.: 0.820Hz (Flap)

§ 3rd Eig.: 1.904Hz (Edge)

§ 4th Eig.: 2.198Hz (Flap)

§ 5th Eig.: 4.045Hz (Possibly coupled)

The first 5 eigenmodes of the AVATAR blade are shown in the figure below.

The flapwise direction is the Y-direction (from suction to pressure side)

Figure 24 Eigenmodes for the blade (1-2 top row, 3-4-5 bottom row)

5.1.5 Buckling Analysis

The first three buckling modes correspond to the same buckling load factor of 1.786 which is

greater than the INNWIND.EU DTU blade (1.75).

These modes are located between between 45.52m – 60.05m on the suction side of the blade

mostly affecting the SPAR CAP elements.

Due to the sensitivity in the buckling analysis (on the spar caps) we have incorporated balsa in

the lamination sequence between the carbon layers. This solution provided results with a

minimum increase on mass and keeping the rest of the properties within required boundaries.

This is not a preferred solution, due to difficulties in manufacturability, but it was used as an

x

y


initial solution. In a production blade there are numerous ways to overcome local buckling that

are better suited to production techniques.

The following two figures illustrate the 1st buckling mode of the blade, which is located between

45.52m – 60.05m. The respective elements of the webs are NOT subject to buckling.

Figure 25 Buckling results (a)

Figure 26 Buckling results (b)

5.1.6 Mass & Stiffness Distribution

The mass/stiffness distribution along the blade is provided in an excel file (AVATAR WP1

participants, 2015) for 51 sections along the blade. Note that the z-axis for the properties

distribution corresponds to the global blade axis from trailing to leading edge, while the y-axis is

x

y

z

x

y

z


again the global blade axis from the compressive side (suction) to the tensile side (pressure) of

the blade.

An example for section at 25.0m is shown with the relevant axis in Figure 27. On this figure also

the elastic centre is shown as an example.

Figure 27 Conventions used in the geometry description

5.2 Redesign of the blade The initial design demonstrated the possibility to construct a lightweight blade at the required

size; however some undesirable properties of the resulting solution needed to be addressed in

order to have a more realistic solution. The identified issues were related to:

Small torsional rigidity. As a result of this there was substantial twist of the blade (3-3.5° at the

tip, compared to 1-1.5° for the INNWIND.EU blade). The increased twist was not a

problem per-se, as it resulted in alleviation of the blade loads, but the resulting power

curve was inferior to the one predicted at the design stage. While performance was

still better than the INNWIND.EU blade, the difference was not as big as expected. It

was therefore decided to increase the blade torsional stiffness in order to rectify this.

Placement of eigenfrequencies. While there was no immediate interaction with the 3P/6P

frequencies, the margin from the 4P frequency was too small for the initial blade. This

resulted in undesirable dynamic behavior, based on the initial aeroelastic simulations.

Large flapwise stiffness. Replacing the glass construction with carbon fiber spars resulted in a

fundamentally different dynamic response of the blade. As a result of the change in

relative stiffness between flap and edge directions, the 1st eigenfrequency now

corresponds to an edge mode with the second one corresponding to a flap mode.

This change is unusual, but not necessarily a problem. The real problem is related to

the fatigue loads calculated during the initial evaluation of the blade. During normal


operation the average bending moments were comparable to the INNWIND.EU blade,

but the spikes from wind speed variations were substantially higher. The reason is

obviously that the smaller deflections of the blade result in larger forces in order to

absorb the same amount of energy.

The combination of these issues led to a decision for a reevaluation of the structural

characteristics of the blade.

Table 5 Basic structural properties of final AVATAR blade

In the following results we make reference to three different blade structural designs

r0 is the initial blade design


r1 is a revision of that with increased stiffness and weight near the hub and increased

torsional stiffness. Edge and flapwise stiffness is unchanged in the outer 75% of the

blade

r2 is a completely changed blade with decreased flap stiffness along the span (and

retaining the increased torsional stiffness)

The first two issues were relatively easy to rectify, using the modified r1 blade. For the third

issue, a simple modification was not sufficient, so a new design was produced starting with the

blade´s desired global structural parameters. The details of mass and stiffness distributions for

the final design are given in Table 5. Full details of the proposed geometry will be available in

the form of a spreadsheet at the project web-site (AVATAR WP1 participants, 2015).

In order to verify that the values used for the new design were acceptable an initial investigation

of the weight increase that comes with it was performed. In order to do that, PoliMI conducted a

redesign of the blade from scratch, in order to deliver a possible internal layout for the AVATAR

blade. In this view, the activity concerns only the structural design of the blade, since the

external shape (chord, twist, airfoils) is kept frozen during the whole design process, in order to

assure consistency with the other activities within the Deliverable. The solution, hereafter

identified by the suffix _r0_PoliMI is briefly illustrated in the following. This solution is only used

to verify that the prescribed mass increase is reasonable. Additional “iterations” would be

needed to satisfy all desired design specifications. It was therefore considered preferable to use

the modified (r2) blade, where the dynamic characteristics have been tuned so as to be in-line

with generally acceptable choices. A more detailed investigation is planned for a later stage of

the project.

5.2.1 Initial Design for weight estimation

The structural redesign of the AVATAR was obtained through the multi-disciplinary optimization

tool Cp-Max (a Code for Performance Maximization), whose main features are extensively

discussed in (Bottasso, et al., 2014) and (Bottasso, Croce, & Campagnolo, 2012). Basically, the

algorithm seeks the minimization of the cost of energy through a multi-level optimization, which

is based on the iterative solution of a closed aero-elastic design loop interfaced with a 3D FEM

analysis, until an optimal solution which satisfies all the constraints is found.

As shown in Figure 28 the ‘coarse-level’ allows the constrained optimization of the rotor blade

by using a 1D geometrically exact beam for aero-servo-elastic multibody analysis and load

calculation, integrated with a 2D FEM cross-sectional model, which is employed for the

estimation of the local mass/stiffness properties and for stress/strain verification. At this stage,

the design is subjected to the following set of local and global constraints:

Prescribed minimum gap between the first flap-wise frequency and the 3P

Prescribed minimum gap between the first edge-wise and the first flap-wise frequencies

Maximum tip deflection

Lower-than-admittable stress/strain in each structural element

Fatigue loads

Preliminary buckling verification


Once the coarse level optimization is completed, the optimal solution is then tested with a ‘fine-

level’ 3D FEM model, in order to identify possible shortcomings of the design, in particular for

what concern the buckling, and to obtain a refined estimation of the non-structural mass.

The improved results obtained at the fine-level are then included in the following coarse-level

iteration through a modification of the constraints.

Figure 28 Multi-level optimization

5.2.1.1 Internal blade layout

The structural model of the AVATAR_r0_PoliMI is based on the typical cross-sectional layout

reported in Figure 29, which illustrates the position of the main elements within the section. In

addition to them, a root reinforcement is defined in the region close to the hub and a third web is

present in the mid part of the blade. Table 6 lists all the 9 structural members which are defined

for this analysis, together with the non-dimensional coordinate η where each member starts and

ends. The two 800mm-wide spar caps start at η=0.05 and their width is held constant along the

entire span. However, as illustrated in Figure , a transitional region is defined between η=0.0

and η=0.05: here the spar is gradually extended in order to embrace the root cylinder and


simultaneously blended with the root reinforcement, in order to ensure a continuous distribution

of material and a smooth transition between the root solution and the rest of the blade.

Each structural element is made up of specific materials, whose properties are taken from

(Lekou, 2014) and reported in Table 7. As specified in the same report, the design of AVATAR

is based on the usage of carbon fibers (CF) in the spar. Here, for consistency, the unidirectional

carbon fiber is assigned also to the root reinforcement, in order to avoid sharp variations of the

structural properties. The design of the other elements, like for example the external shell, the

webs and the LE-TE reinforcements is based on glass-fibers composites (GF).

Figure 29 Cross-sectional arrangement of the structural elements

Figure 30 Planform view of the blade


ID Element ηStart ηEnd Material(s)

1 Shell 0.0 1.0 Triaxial, Balsa 2 Suction side spar cap 0.05 0.9812 Unidirectional (CF) 3 Aft shear web 0.05 0.9812 Biaxial, Balsa 4 Pressure side spar cap 0.05 0.9812 Unidirectional (CF) 5 Rear shear web 0.05 0.9812 Biaxial, Balsa 6 LE reinforcement 0.15 0.95 Unidirectional (GF), Balsa 7 TE reinforcement 0.15 0.95 Unidirectional (GF), Balsa 8 Root reinforcement 0.0 0.65 Unidirectional (CF) 9 Third web 0.22 0.95 Triaxial

Table 6 Structural members

TRIAXIAL UNIAXIAL (CF) UNIAXIAL (GF) BIAXIAL

E11 [GPa] 21.790 115.0 41.630 13.920

E22 [GPa] 14.670 7.560 14.930 13.920

ν12 [-] 0.478 0.300 0.241 0.533

G12 [GPa] 9.413 3.960 5.047 11.500

ρ [Kg/m3

] 1845 1578 1915 1845 σ

11_Ten [MPa] 480.4 1317.60 876.1 223.2

σ11_Comp

[MPa] 393.0 620.13 625.8 209.2 σ

22_Ten [MPa] 90.4 21.88 74.03 223.2

σ22_Comp

[MPa] 152.7 76.25 189.4 209.2 τ

12 [MPa] 114.0 45.53 56.58 140.3

Table 7 Properties of the composite materials

In addition to the design of the structural elements, several contributions to the non structural

mass are taken into account. These are summarized in Table 8 and Table 9.

BALSA PAINT ADHESIVE

E [Pa] 0.05e+09 1.00 4.56e+06

G [Pa] 0.15e+09 0.384 1.45e+06

ρ [Kg/m3

] 110 1400 1100

σ11_Ten

[MPa] 0.694 10.0 61.5

σ11_Comp

[MPa] 0.4 10.0 65.3

τ12 [MPa] 0.307 10.0 36.6

Table 8 Properties of balsa, paint and adhesive

Type Element Units Value

Linear NSM


Glue mass shear web [kg/m] 2.30 Glue mass LE, TE [kg/m] 2.88 Lightning system [kg/m] 0.50 Bonding plies shear web [kg/m] 3.90 Surface NSM Resin uptake shell [kg/m^2] 2.50 Resin uptake shear web [kg/m^2] 0.80

Table 9 Non-structural mass

5.2.1.2 Optimal design

After several iterations, the optimization provided the thickness distributions illustrated in Figure

31. As expected, the design of the spar caps was the most labour-intensive aspect of the

procedure, since it required several adjustments in order to avoid buckling.

Figure 31 AVATAR_r0_PoliMI: Thickness of the structural members

The main structural properties of the blade are compared against the AVATAR_r2 in Figure 32

and Figure 33.


Figure 32 AVATAR_r0_PoliMI: Flap-wise (up) and edge-wise (down) stiffness distribution

Figure 33 AVATAR_r0_PoliMI: mass per unit length distribution

5.2.1.3 Constraints

The following figures illustrate the value of the constraints for each structural element along the

blade. It must be recalled that a negative value of the bars implies that the constraint is satisfied.


Shell

Suction side spar cap

Forward shear web

LE - reinforcement

TE - reinforcement

Root reinforcement

Figure 34 AVATAR_r0_PoliMI: Value of the constraints along the blade

In addition, Table 10 lists, for each element, the most active constraint which affects the design.


ID Element Most active design constraint

1 Shell Fatigue 2 Suction side spar cap Buckling 3 Forward shear web Fatigue 4 Pressure side spar cap Buckling 5 Rear shear web Fatigue 6 LE reinforcement Frequencies separation 7 TE reinforcement Frequencies separation 8 Root reinforcement Stress/strain 9 Third web None

Table 10 AVATAR_r0_PoliMI: Active constraints

5.2.1.4 Tip deflection

The maximum tip deflection for the AVATAR_r0_PoliMI was d=8.86 m, and occurred on blade 3

during the DLC 2.3 (EOG plus occurrence of fault), simulated at 25 m/s, which implies that the

maximum tip deflection is not an active constraint for the design.

5.2.1.5 Eigenfrequencies

The following clamped-blade eigenfrequencies were identified:

Mode Frequency [Hz] Type

1 0.665 Flap 2 0.777 Edge 3 1.91 Flap 4 2.31 Edge 5 4.04 Flap

Table 11 AVATAR_r0_PoliMI: First 5 eigenfrequencies of the clamped blade

And the corresponding modal shapes are illustrated in Figure 35. Considering an operating

rotational speed equal to 9.6 rpm, the first flapwise frequency is 38% higher than the 3P. At the

same time, the first edgewise frequency is 16.8% higher than the first flapwise.


1. First flap-wise mode at f=0.665 Hz

2. First edge-wise mode at f=0.777 Hz

3. Second flap-wise mode at f=1.91 Hz

4. Second edge-wise mode at f=2.31 Hz

Figure 35 AVATAR_r0_PoliMI: modes of the clamped blade

5.2.1.6 Mass

The total blade mass of the AVATAR_r0_PoliMI is 48118 kg. Table 12 shows the contribution of

each structural element to the total mass, as well as the contributions of the various non-

structural masses.


Element Mass [kg] % of total mass

Spar caps 16439 34.16

Root reinforcement 6628 13.77

Shell 6901 14.34

LE - TE reinforcements 4854 10.09

Shear webs 2866 5.96

Third web 110 0.23

Total structural mass: 37798 78.55

Surface NSM 2180 4.53

Linear NSM 1980 4.11

Paint 735 1.53

Core (webs) 1619 3.36

Core (shell) 2429 5.05

Core (LE-TE reinforcements) 1377 2.86

Total NSM: 10320 21.45

Total blade mass: 48118 100.0

Table 12 AVATAR_r0_PoliMI: mass contributions from individual elements

5.2.1.7 Loads envelope

The DLC considered during the design of the AVATAR_r0_PoliMI are listed in Table 13Table .

These cases were chosen in order to include the worst-case scenario for what concerns both

the maximum tip displacement and the loads.

Case DLC Wind model Number of simulations Wind speed

Power production

1.1 NTM 12 Vcutin ≤ V ≤ Vcutout 1.3 ETM 12 Vcutin ≤ V ≤ Vcutout

1.4 ECD 16 Vr-2 ≤ V ≤ Vr+2 Production plus fault

2.3 EOG Vr-2 ≤ V ≤ Vcutout

Parked 6.2 EWM-50 12 50 m/s

Table 13 Considered DLC

The loads envelope at the blade root for the three blades is reported in Table 14. The

corresponding reference system is sketched in Figure 36.


Figure 36 Blade coordinate system

In-plane

Bending Mx

Out-of-plane

Bending My

Combined

Bending Mxy

Torsion

Mz Blade DLC

[kNm] [kNm] [kNm] [kNm] - -

Max 65441 -54245 85000 255 2 DLC23_voa

Min -50310 -2768 50386 -165 3 DLC23_voa

Max -4026 81766 81865 -80 2 DLC11_25a

Min 39226 -80679 89709 206 3 DLC23_voa

Max 39226 -80679 89709 206 3 DLC23_voa

Max 59214 -63959 87161 466 3 DLC23_vr+2a

Min 2434 -29475 29575 -902 3 DLC14_vrplusa

Table 14 AVATAR_r0_PoliMI: Loads envelope at blade root for the three blades

The loads envelope at the tower top is reported in Table 15 following the coordinate system

sketched in Figure 37, while the load envelope for the hub is reported in Table 16, following the

coordinate system of Figure 38


Figure 37 tower top coordinate system

Side-side

Bending Mx

Fore-aft

Bending My

Combined

Bending Mxy

Torsion

Mz DLC

[kNm] [kNm] [kNm] [kNm] -

Max 27915 6170 28589 17191

DLC62_ID_T_YMdeg-

90a

Min -13186 3326 13599 -9662

DLC62_ID_T_YMdeg90

a

Max 13596 93874 94853 -28881 DLC11_25a

Min -462 -83445 83446 7604 DLC23_vr+2a

Max 13596 93874 94853 -28881 DLC11_25a

Max 19230 3487 19543 46616 DLC11_25a

Min -4291 -35017 35279 -68335 DLC23_voa

Table 15 AVATAR_r0_PoliMI: Loads envelope at tower top


Figure 38 Hub coordinate system

Fx My Mz Myz DLC

[kN] [kNm] [kNm] [kNm] -

Max 3212 -35790 -10723 37362 DLC23_voa

Min -3621 57118 -2315 57165 DLC23_vr+2a

Max -3594 58613 7733 59121 DLC23_vr+2a

Min 998 -95442 -21124 97752 DLC11_25a

Max 2067 -5556 44989 45330 DLC11_25a

Min -2275 21071 -66248 69518 DLC23_voa

Max 1078 -95140 -27667 99082 DLC11_25a

Table 16 AVATAR_r0_PoliMI: Loads envelope at hub


6 Loads and Stability

6.1 Test Cases For the initial evaluation, a limited set of test cases were used, while a more complete set will be

used for the final evaluation. The tests that were performed included

Steady operation for rigid and flexible blades, in order to verify the design power curve of

the wind turbine

Stability analysis, including uncoupled and coupled eigen-frequencies of the blades

Unsteady simulations of the operation during speed-up and speed-down, in order to

identify obvious problems and verify the suitability of the controller used for the wind

turbine

Unsteady operation in normal conditions (DLC1.2), in order to get a first estimate of the

expected fatigue loads and displacements in normal conditions

Some partners performed additional tests, based on other Design Load Cases, as specified by

the IEC standards. The tests described in this section have been performed by all three

partners, though in some cases only indicative results from one or two are presented, as the

results were similar. A much more detailed analysis including a full set of DLC cases is to be

reported in D1.3, where comparisons between the results are made.

6.2 Computational Tools Three different computational methods were independently used to assess the aeroelastic

performance of the reference blade.

DTU made use of HAWC2 (Horizontal Axis Wind turbine simulation Code 2nd generation) which

is an aeroelastic code intended for calculating wind turbine response in time domain.

The core of the code was developed mainly within the years 2003-2007, by the Aeroelastic

Design Research Program at DTU Wind Energy, DTU Risø Campus in Denmark. HAWC2 is

developed and distributed by DTU Wind Energy and has been used in numerous research

projects and industrial applications. HAWC2 has a large number of users and is used both for

design and verification purposes.

HAWC2 is able to simulate wind turbines in time response with following properties:

Normal onshore with 1,2, 3 or more blades

Pitch and (active) stall controlled wind turbines

Guyed support structures

Offshore turbines on monopoles, tripods or jackets

Floating turbines with mooring lines

Multiple rotors in one simulation

Multibody formulation that can handle multiple degrees of freedom (like blade torsion)

Detailed aerodynamic model that includes:


o Dynamic stall models: Stig Øye model, a modified Beddoes-leishmann model

and a model for ATEF (Active Trailing Edge Flaps)

o Skew inflow model

o Shear effects on the induction

o Dynamic inflow model

Hydrodynamic model based on Morrison’s equation

Water Kinematics that includes:

o Currents

o Linear airy waves

o Irregular airy waves

o Deterministic irregular waves

o Stream function wave

Wind, turbulence and wake models:

Control interface performed through DLL´s (Dynamic Link Library)

Default controller provided with a pitchregulated variable speed controller

Eigenvalue analysis at standstill

PoliMi made use of the software Cp-Lambda (Code for Performance, Loads and Aeroelasticity

by Multi-Body Dynamic Analysis), based on a finite-element multibody formulation.

The multi-body approach is based on the full finite-element method, i.e., no modal-based

reduction is performed on the deformable components of the structure. Cartesian coordinates

are used for the description of all entities in the model, and all degrees of freedom are referred

to a single inertial frame; the formulation handles arbitrarily large three-dimensional rotations.

The turbine blades, the tower and the shaft are modeled using geometrically exact, composite-

ready beams. The formulation models beams of arbitrary geometry, including curved and

twisted reference lines, and accounts for axial, shear, bending, and torsional stiffness. Joints are

modeled through holonomic or nonholonomic constraints, as appropriate, that are enforced by

means of Lagrange multipliers using the scaled augmented Lagrangian method. All joints can

be equipped with internal springs, dampers, backlash, friction, and power loss models, which

are used among other things to account for such effects in the gear-box and drive-train.

The blade pitch system response is modeled with a second order system, while the response of

the generator by a first order one; both actuator models receive commanded signals by the

supervision and feedback controllers.

Lifting lines can be associated with beam elements and their geometric description is given in

terms of three-dimensional twisted curves; for generality of the implementation, these

aerodynamic reference curves are distinct from the structural reference ones they are

associated with. The lifting lines are based on classical two-dimensional blade element theory,

and account for the aerodynamic center offset, twist, sweep, and unsteady corrections. At a

number of span-wise stations along each lifting line, the aerodynamic characteristics of the

airfoil used at that location are given using look-up tables, which store for a given number of

angles of attack and Reynolds numbers the values of the sectional lift, drag, and moment

coefficients. Lifting lines are used here to model the aerodynamic characteristics of the blades,


but also of the tower and of the nacelle. An inflow element can be associated with the blade

lifting lines so as to model the rotor inflow effects; the code implements the Peters–He dynamic

inflow wake model and a classical blade-element momentum (BEM) model based on the

annular stream-tube theory with wake swirl. Tip and hub loss models are also considered.

Wind is modeled as the sum of a steady state mean wind and a perturbation wind, accounting

for turbulence and/or gusts. The deterministic component of the wind field implements the

transients (i.e. gusts), the exponential and logarithmic wind shear models, and the tower

shadow effects, which include the potential flow model for a conical tower, the downwind

empirical model or an interpolation of these two models. The stochastic component of the wind

field is computed according to the Mann or Kaimal turbulence models. The turbulent wind is pre-

computed before the beginning of the simulation for an assigned duration of time and for a user-

specified two-dimensional grid of points.

During the simulation, the current position of each airstation is mapped to this grid, and the

current value of the wind is interpolated in space and time from the saved data. The multi-body

formulation used in this effort leads to a set of non-linear partial differential algebraic equations.

Spatial discretization of the flexible elements of the model using the finite-element method yields

a system of differential algebraic equations in time that are solved using an implicit integration

procedure that is nonlinearly unconditionally stable. The implicit nature of the scheme allows for

the use of large time steps and is more appropriate than explicit schemes for the typical

dynamics of rotor systems. At each time step, the resulting non-linear system of equations is

solved using a quasi-Newton scheme. The time-step length is adjusted based on an error

indicator. The code supports static and transient analyses, and the computation of

eigenfrequencies and eigenmodes about deformed equilibrium configurations.

CENER has used GL GH BLADED v4.4 for the calculations into D1.2. GL GH BLADED is a

BEM based dynamic load calculation program for multibody structures with coupled

aerodynamic and hydrodynamic loads. BLADED calculates the structure dynamic with a

multibody-dynamic-approach. The components such as blades and tower are modelled from

single, linear and flexible elements whose deformation is determined by modal analyses. This is

done by a linear combination of the calculated eigenmodes resulting from a finite elements

method calculation. The aerodynamic forces are calculated with the blade element momentum

theory. In addition, BLADED is able to simulate a wake behind the rotor and dynamic stall.

6.3 Results

6.3.1 Power Curve Verification

For these tests the operation of the blade under constant (non-dynamic) conditions was

examined for stiff (theoretical) and flexible (actual) blade. For the stiff blade, the results closely

matched the predicted design performance (as expected, since they are both based on BEM

modeling). For the flexible blade there was a reduction in power production, as a result of the

blade twist. As discussed in the previous section, this was initially quite large, but for the

redesigned (r2) blade, with increased torsional stiffness, the performance was improved. In all

cases there was an improvement in energy yield compared to the INNWIND.EU blade. A

detailed investigation of the calculated power curve is performed as part of WP2.


Figure 39 Comparison of power curves for the various configurations and comparison with INNWIND.EU rotor stability analysis

The first test performed was the estimation of the blade natural frequencies, without taking into

account rotation effects. The results from this are summarized in Table 17, showing a generally

good agreement between results obtained with different calculation methods, and a substantial

separation of the prevailing eigenfrequencies.

Table 17 Blade eigenfrequencies

CENER PoliMI

1st flap 0.684 (4.27P) 0.6733 (4.21P) 1st edge 0.901 (5.63P) 0.8861 (5.54P) 2nd flap 1.883 (11.77P) 1.8335 (11.46P) 2nd edge 2.606 (16.29P) 2.4912 (15.57P) 5th 3.852 (24.05P) 3.6884 (23.05P)

The second step in the stability analysis was the derivation of the Campbell diagrams for the

operating envelope of the proposed wind turbine. Again, no serious problems were identified

with the final design.


Figure 40 AVATAR RWT Campbell diagram (CENER)

Figure 41 AVATAR RWT Campbell diagram (PoliMi)


Figure 42 AVATAR RWT Campbell diagram (DTU)

6.3.2 Speed-up/Speed-down simulations

A ramp of wind speeds from 0 to 25m/s was used to test the behavior of the controller and

ensure that there was not an undesirable behavior in transitional regions. The results show that

there are some excursions during the transitional period, but not at alarming levels. Given that

the controller has not been optimized for this wind turbine, the results are considered

satisfactory. A check of the blade deflections and torsion at tip (Figure 48) shows that the

redesigned blade displays reduced torsion and increased flapwise deflections, as intended,

compared to the earlier designs.

Figure 43 Simulation of step-up operation (CENER)


Figure 44 Simulation of step-down operation (CENER)

Figure 45 Simulation of step-up operation (PoliMi)

Figure 46 Simulation of step-down operation (PoliMi)


Figure 47 Simulation of step-up and step-down operation (DTU)

Figure 48 Changes in torsion and deflection between the three designs (PoliMi)

6.3.3 Unsteady operation

For the unsteady operation, a series of runs for different wind speeds were performed by the

participating partners.


Figure 49 Displacements for DLC1.2 at different wind speeds (CENER)

Figure 50 Bending moments for DLC1.2 at different wind speeds (CENER)


Figure 51 Deflection and bending moment at 11m/s (DTU)

As seen in e.g. Table 18. The proposed blade is still stiffer than the equivalent INNWIND.EU

blade, resulting in smaller deflections (out-of-plane), while the deflections in-plane are slightly

increased as a result of the increased length and weight.

Table 18 Max calculated values of deflection and loads during normal operation (DLC1.2) (CENER)

Blade Tip Deflection (Out of Plane) 8.05m Blade Tip Deflection (In Plane) 2.23m Blade Tip Torsion (degs) 2.52 Blade Root Bending Moment (Out of Plane) 50432kNm Blade Root Bending Moment (Out of Plane) 21311kNm

Table 19 Max. loads based on all DLC cases used (PoliMi)


Finally, an estimate of the fatigue, based on the DLC12 runs was performed. This is not an

exhaustive fatigue analysis, but an effort to estimate the relative level of loads compared to the

established INNWIND.EU blade. A summary of the results is given in Table 20. An increase of

~30% in equivalent loads is expected, which is in line with the additional length and weight of

the blade.

Table 20 Damage Equivalent loads estimate (PoliMi)


7 Conclusions The initial design process and the resulting blade and wind turbine parameters for a reference

high speed wind turbine have been presented in the report. The resulting values will be used for

the work performed for the other work packages of the AVATAR project.

The main drivers for the design were to produce a blade that would

a) Stretch the limits of the computational methods currently in use

b) Have a performance benefit compared to the INNWIND.EU blade, on which it is based (so

that there is justification for the proposed design direction)

c) Provide a reasonable basis for further design changes in the following stages of the project.

The current blade is not necessarily the best solution for the problem in hand. As already

discussed, there are different ways to achieve the lower power density that is required for the

low induction rotor. The best way is probably to use low-lift airfoils, e.g. airfoils having their

maximum k = CL/CD at moderate CL values. These will be evaluated at a later stage, as it was

necessary to work with existing airfoil families for the reference design. Alternatively, part of the

required reduction in 𝑐 ∙ 𝐶𝐿 could come from a reduction in chord. Smaller chords allow for better

aerodynamic performance (higher k values) and local buckling resistance but lead to heavier

and softer designs (using the same materials). Such a solution would lead to an increase in tip

deflection and would change significantly the dynamics of a large blade getting, in this case, its

first natural frequency down from 4P close to 3P.

Changing the construction materials (from glass fibre to carbon spars) can alleviate this

problem, leading to significantly stiffer (though more expensive) blades. Evidently, to get the

best compromise, the design space needs to be checked further for the 2nd stage optimum

design, in order to get the optimum combination of lift, chord and blade thickness for very large

blades.


8 References AVATAR WP1 participants. (2015, 2). AVATAR10MWReferenceWindTurbine.xls. Retrieved

from AVATAR Intranet: http://www.eera-avatar.eu

Bak, C., Zahle, F., Bitsche, R., Kim, T., Yde, A., Henriksen, L., . . . A. Natarajan, M. H. (n.d.).

Design and performance of a 10 MW wind turbine (to be accepted). J. Wind Energy.

Bottasso, C., Campagnolo, F., Croce, A., Dilli, S., Gualdoni, F., & Nielsen, M. (2014). Structural

optimization of wind turbine rotor blades by multi-level/sectional/multibody/3D FEM

analyisis. Multibody System Dynamics, Vol. 32, 87-116.

Bottasso, C., Croce, A., & Campagnolo, F. (2012). Multi-disciplinary constrained optimization of

wind turbines. Multibody System Dynamics Vol. 27, 21-53.

Chaviaropoulos, T., Beurskens H.J, M., & Voutsinas, S. (2013). Moving towards large(r) high

speed rotors – is that a good idea? EWEA Conference Proceedings. Vienna.

Griffith, D. T. (2013). The SNL-100-01 Blade: Carbod Design Studies for the Sandia 100m

Blade.

Griffith, D. T., & Ashwill, T. D. (2011). The Sandia 100-meter all-glass baseline wind turbine

blade: SNL100-00. New Mexico: Sandia National Laboratories.

IEC. (2005, August). IEC 61400-1: Wind Turbines - Part 1: Design Requirements. Standard.

Standard.

Lekou, D. (2014). Report on structural solution for a long carbon blade within AVATAR project.

CRES, Centre for Renewable Energy Sources and saving.

Stettner, M., & Chaviaropoulos, P. (2014). Reference Blade Specifications. Deliverable D1.1.

Documents

AVATAR Reference Blade Design