Journal cover Journal topic
Wind Energy Science The interactive open-access journal of the European Academy of Wind Energy
Journal topic
Wind Energ. Sci., 3, 163-172, 2018
https://doi.org/10.5194/wes-3-163-2018

Special issue: Wind Energy Science Conference 2017

Wind Energ. Sci., 3, 163-172, 2018
https://doi.org/10.5194/wes-3-163-2018

Research articles 03 Apr 2018

Research articles | 03 Apr 2018

# Benefits of subcomponent over full-scale blade testing elaborated on a trailing-edge bond line design validation

Malo Rosemeier, Gregor Basters, and Alexandros Antoniou Malo Rosemeier et al.
• Division Structural Components, Fraunhofer IWES, Fraunhofer Institute for Wind Energy Systems, Am Seedeich 45, 27572 Bremerhaven, Germany
Abstract

1 Introduction

Although (Det Norske Veritas Germanischer Lloyd Aksjeselskap) accepts combined static full-scale tests (SFST) rather than pure lead–lag and flapwise, the importance of which is highlighted in and , this approach is rarely used for certification tests.

Attempts have been made to improve fatigue full-scale blade testing (FFST) towards a more realistic scenario by minimizing the overloading, i.e., the deviation between applied test bending moments and target bending moments as per the design requirements. For example, applied an algorithm to optimize the bending moment distribution for uni-axial, resonant FFST by using additional masses and determining optimal actuator positions and excitation frequencies. Moreover, propose a uni-axial multifrequency approach to replicate the actual spatial damage distribution more realistically. Bi-axial FFST is an approach that emulates more comprehensive damage along the blade's circumference . The random-phase bi-axial resonant excitation has been a focus of research for a number of years . Yet another approach is the phase-locked excitation with frequency ratios, e.g., of 1 : 1 between the flapwise and lead–lag mode . The flapwise mode is excited in its eigenfrequency while the lead–lag mode is forced in the same frequency . Alternatively, both modes can be excited in resonance by tuning both eigenfrequencies independently with the concept of virtual masses (Post2016).

IEC (2012) states that during fatigue testing, the mean loads applied shall normally be as close as possible to the mean load under the operating conditions that cause the most severe fatigue damage. The importance of considering mean loading for the fatigue life evaluation of fiber-reinforced polymers is described and quantified in .

, and proposed subcomponent test (SCT) concepts to investigate the structural performance of trailing-edge bond lines in particular. showed an alternative SCT method for investigating the flapwise buckling response of a blade segment.

Within this work, SCT is presented as a means to potentially overcome some of the drawbacks of FST. The intention of applying SCT is not to overcome the necessity of FST. The method of SCT, however, has the potential to be considered as a standard intermediate test before, during or after certification. By taking the target mean loading and thus the stress ratio between the minimum and the maximum stress exposure at the trailing edge into account, the structural evaluation of the blade can be approached more realistically compared to field simulations. The testing time can be reduced significantly by means of SCT when considering the case of a particular blade segment validation, recertification due to a repair or design retrofit solution.

First, the FST and SCT concepts are described. Second, on the basis of calculations the benefits of SCT over static and fatigue FST are highlighted.

2 Full-scale blade and subcomponent testing concepts

For a SFST, the blade is clamped to a stiff block and several load frames are attached to the blade along the span1 (Fig. 1a). Point loads are introduced via load frames pulling ropes which are connected to actuators on a strong floor or strong wall.

For a uni-axial, resonant FFST, only one load frame is attached to the blade. Cyclic tension and compression loads are applied by one actuator (Fig. 1b). The testing frequency is chosen so as to be close to the blade's eigenfrequency, thus enabling displacement-controlled resonant excitation of the mode shape of the blade.

The concept of blade SCT discussed in this work is based on the work by , where a subcomponent is defined as a part or section representative of the blade design (IEC2017). The specimen can either be a full-scale blade cutout or a separately manufactured full-scale part. In the present study the area of interest is the trailing-edge cell, including one main shear web (Fig. 1c), to obtain a closed cross section, which can still realistically emulate “breathing” or “pumping” effects of the thin-walled blade structure. Both ends of the subcomponent are glued into load frames, which are connected to ball joints. An eccentric axial load is introduced at one joint, thus introducing a normal force on which a bending moment is superimposed. The position of the two joints within the cross sectional plane can be chosen arbitrarily, which makes it possible to introduce any load combination and distribution of lead–lag and flapwise loading.

3 Comparison of subcomponent with full-scale blade testing

In the following, the benefits and drawbacks of SCT over static and fatigue full-scale blade testing (FST) are elaborated.

An advantage of SCT over FST is that the number of relatively inexpensive specimens can be increased to investigate different design variants of a critical blade detail, for example. Owing to the smaller dimension and relatively small displacements of an SCT, it is possible to conduct experiments under environmental conditions, e.g., in climate chambers, and to better observe the experiment via optical measurement systems, such as the digital image correlation shown by .

A disadvantage of SCT is that only a segment of the blade is considered for testing. Furthermore, the design of the boundary conditions of an SCT relies on detailed models, e.g., for the determination of blade properties of a cutout segment as explained in .

The focus of the following sections is on the problem of constrained areas and overloading of blade parts, the load direction in static and fatigue testing, as well as on the stress ratio at the trailing edge and the resulting testing time.

Figure 1Load deviation ΔM between test bending moment and target bending moment determined for the DTU10MW blade and test setup for static full-scale test (SFST) in side view (excerpt from (a), fatigue full-scale test (FFST) in top view (b) and blade cutout subcomponent test (SCT) (c). Untested areas are shown in gray.

Load frames diminish tested areas because they constrain the structure (Fig. 1a). The longitudinal dimension of the constrained area at a load clamp is assumed to be ±75 % of the local blade chord length (IEC2012). Critical cross sections should not be within constrained areas. Owing to the multilinear moment distribution in the static test, the critical areas are necessarily in under- or overloaded areas (Fig. 1a). Thus, due to the setup of an FST, these areas are subjected to a load deviation, which means a lower or a higher test loading Mtest compared to the target load envelope Mtarget:

$\begin{array}{}\text{(1)}& \mathrm{\Delta }M=\frac{{M}_{\mathrm{test}}}{{M}_{\mathrm{target}}}.\end{array}$

In the fatigue FST, the constrained area is reduced to solely one load clamp, which increases the area for this test (Fig. 1b). The load deviation ΔM, however, can be even higher for the blade (overloading by up to 20 %) although mass tuning is conducted. Overloading occurs along all tested areas because the shape of the target moment distribution cannot be fully replicated by the test moment distribution. That is, parts of the blade can be damaged far beyond their fatigue life before all relevant areas reach their target damage.

In an SCT, the constrained areas are determined by means of a method similar to that used in FST (Fig. 1c). This implies that a reasonable length of the specimen should be chosen to replicate, for example, a realistic buckling response. The overloading in the tested areas, however, is close to zero, since the bending moment can be adjusted by specially designed boundary conditions .

Moreover, in an SCT the cross section of interest can be tested up to its final fatigue limit state compared to an FST, where other, highly overloaded blade areas may limit the damage that is required to reach the final fatigue limit state at the cross section of interest.

Figure 2Load envelopes in terms of moment vectors determined from aeroservoelastic simulations for several cross sections of a 34 m blade and certification test loads contrasted with strength and stability resistance of the whole blade's cross sections from 0 to 100 % blade length (lB) in terms of the root bending moment (excerpt from ). The two moment vectors represent the critical load directions at 35 and 50 % lB.

## 3.2 Load direction in static testing

In Fig. 2 the load envelopes at different cross sections along the blade are plotted against the strength and stability resistance envelope of all cross sections, i.e., the whole blade. Furthermore, the load direction with the critical stress exposure is highlighted by way of example for two different cross sections at 35 % blade length (lB) and 50 % lB, where the critical load directions are a combination of PTS plus LTT and STP plus LTT, respectively.

Theoretically, the load envelopes should be compared to the resistance envelopes at each particular cross section to determine the most critical load directions. The determination of the resistances for each cross section is only possible via analysis of a blade segment or subcomponent cutout from the full blade subjected to the load envelope determined from aeroservoelastic simulations.

## 3.3 Load direction in fatigue testing

With regard to the load direction, the uni-axial, resonant FFST is constrained to the static blade dead load (suction-side points towards the strong floor) superimposed with the dynamic inertia loads of the mode shape of the blade. According to IEC (2012), the mean loads applied during FFST shall be as close as possible to the mean load at the operating conditions that cause the most severe fatigue damage.

Since no load directions other than the first and second modal shapes, i.e., lead–lag and flap, are possible, the only degree of freedom during uni-axial, resonant FST is the direction of the gravity load due to the pitching position of the blade.

The loads for FFST and field calculations are elaborated by way of example using the DTU10MW reference turbine with respect to the loading of the trailing-edge bond line. A simplified load calculation was conducted comprising the superposition of the quasistatic mean flapwise aerodynamic load for 10 wind speed bins plus the lead–lag gravity load due to the rotor revolution as a function of the blade pitch angle.

A beam model implemented in ANSYS APDL (ANalysis SYStem Parametric Design Language) (Swanson2014) was assembled with a fully populated cross section stiffness matrix determined by the BEam Cross Section Analysis Software, BECAS . The blade parameterization and input generation were conducted using workflows from the FUSED-Wind framework . The BEM-based aerodynamic rotor simulator CCBlade (Ning2014) was used and populated with airfoil polars determined by Rfoil, an extension of Xfoil (Drela1989) including rotational effects .

The lead–lag load cycles were determined on the basis of wind speed distribution and the rotor revolution as well as the design life of 20 years. Furthermore, damage-equivalent loads were determined following using an S–N exponent of m=10, which is a typical value for a glass-fiber-reinforced epoxy (GL2010) and ultimate loads as of .

The greatest damage impact was found at wind speeds around conditions rated between 11.0 and 15.6 m s−1 (Figs. 3 and 4). found similar results for a 34 m blade. The relative damage impact peak at around the rated wind speed of an outboard cross section (Fig. 4) protrudes over the peak of an inboard cross section (Fig. 3) because in outboard regions the influence of aerodynamic loads predominates.

Figure 3Relative damage of different wind speed bins at 10 % blade length.

Figure 4Relative damage of different wind speed bins at 70 % blade length.

Figure 5Moment vectors of lead–lag loading at different wind speed bins and in a lead–lag fatigue full-scale blade test (LL-FFST) at different pitch angles at 10 % blade length. A flapwise FFST (F-FFST) and the bounding box of possible bi-axial FFST scenarios (B-FFST) are shown.

Figure 6Moment vectors of lead–lag loading at different wind speed bins and in a lead–lag fatigue full-scale blade test (LL-FFST) at different pitch angles at 70 % blade length. A flapwise FFST (F-FFST) and the bounding box of possible bi-axial FFST scenarios (B-FFST) are shown.

Furthermore, uni-axial, resonant, flapwise and lead–lag FFSTs were designed. To this end, the flapwise and lead–lag target loads were calculated on the basis of damage-equivalent loads and scaled according to and to cycle numbers of respectively 1.0×106 and 3.0×106 using an S–N exponent of m=10. The test loads were determined on the basis of the inertia loads caused by the dynamic excitation of the first flapwise and lead–lag dominated eigenmode extracted from a modal analysis. This load distribution was superimposed with the gravity loading according to the FFST setup for different pitch angles θ, where $\mathit{\theta }=\mathrm{0}{}^{\circ }$ corresponds to the suction side pointing towards the strong floor and a negative pitch angle which rotates the leading edge towards the floor.

The minimum and maximum bending moment vectors acting at each wind speed bin according to field calculations and for three pitched lead–lag FFSTs are shown for two cross sections at 10 and 70 % blade length (Figs. 5 and 6). From the angle between the minimum and maximum moment vectors at the inboard cross section it can be seen that the lead–lag gravity loads dominate. Moreover, bending moment vectors for the flapwise FFST are superimposed with the lead–lag FFST. The resulting bounding box represents the extremes of the bending moment vector in different bi-axial testing scenarios, i.e., random-phase and phase-locked excitation with different flapwise and lead–lag testing frequency ratios (Post2016). The minimum and maximum bending moment vectors are highlighted for a bi-axial testing scenario that is closest to the field conditions.

Figure 7Axial strain along the trailing-edge bond line at different wind speed bins, lead–lag (LL-FFST) and bi-axial (B-FFST) fatigue full-scale blade tests at different pitch angles.

Using the loading of the most severe wind speeds (11.0, 13.3, 15.7 m s−1), the longitudinal strains in the direction of the blade span at the trailing-edge bond line were determined for the minimum and maximum amplitudes (Fig. 7). It can be seen that the trailing-edge bond line is loaded more in compression during an FFST at $\mathit{\theta }=\mathrm{0}{}^{\circ }$ compared to the strains determined from field load calculations.

Assuming linear elastic material behavior, the stress ratio R can be expressed as the strain ratio between the minimum and maximum longitudinal strains:

$\begin{array}{}\text{(2)}& R=\frac{{\mathit{\epsilon }}_{\mathrm{lmin}}}{{\mathit{\epsilon }}_{\mathrm{lmax}}}.\end{array}$

Considering the R of the field load calculation (Fig. 8), the deviation becomes even more prominent towards the tip, where the loading tends more towards a tension–tension than a tension–compression loading condition. Pitching the blade in an FFST increases the stress ratio, but the inclination of the stress ratio distribution is the opposite to that obtained in field simulations. Thus, in an FFST only the stress ratio of a defined cross section can be obtained via pitching. As already stated above, in an SCT the stress ratio can be adjusted according to the cross section of interest.

Figure 8Stress ratio R of the axial strain along the trailing-edge bond line at different wind speed bins, and lead–lag (LL-FFST) and bi-axial (B-FFST) fatigue full-scale blade tests at different pitch angles.

## 3.4 Stress ratio and testing time

The testing time is compared for the two testing concepts FFST and SCT. To this end, the loading at the trailing-edge bond line due to FFST is compared to the field calculations.

On the basis of the formulation by Puck (1996), the internal loading of a material is expressed in terms of stress exposure e (also referred to as effort), which means an ambient stress σ over the allowable stress (fracture resistance), which is here the tensile strength Rt. Thus the stress exposure is generally defined as

$\begin{array}{}\text{(3)}& e=\frac{\mathit{\sigma }}{{R}^{\mathrm{t}}}.\end{array}$

Furthermore, the materials of a rotor blade, i.e., adhesive, resin and glass fiber, can be considered to be isotropic. Assuming that a symmetric constant life diagram is applicable for an isotropic material , the allowable cycle number to failure Ni for a given load collective i, with a mean stress exposure ${e}_{i\mathrm{m}}=\frac{|{\mathit{\sigma }}_{i\mathrm{m}}|}{{R}^{\mathrm{t}}}$ and a stress exposure amplitude ${e}_{i\mathrm{a}}=\frac{|{\mathit{\sigma }}_{i\mathrm{a}}|}{{R}^{\mathrm{t}}}$, is derived from

$\begin{array}{}\text{(4)}& {N}_{i}={\left(\frac{\mathrm{1}-{e}_{i\mathrm{m}}}{{e}_{i\mathrm{a}}}\right)}^{m}.\end{array}$

This means that the cycle number is directly related to the mean stress exposure, the stress exposure amplitude and to the material-dependent S–N curve exponent m.

Assuming the same S–N curve exponent m for different stress ratios,

$\begin{array}{}\text{(5)}& R=\frac{{e}_{i\mathrm{min}}}{{e}_{i\mathrm{max}}}=\frac{{e}_{i\mathrm{m}}-{e}_{i\mathrm{a}}}{{e}_{i\mathrm{m}}+{e}_{i\mathrm{a}}},\end{array}$

the allowable load cycle number of a load collective of any ratio NiR and the allowable load cycles ${N}_{iR=-\mathrm{1}}$ can be expressed as the relation (Appendix A):

$\begin{array}{}\text{(6)}& \frac{{N}_{iR}}{{N}_{iR=-\mathrm{1}}}={\left(\mathrm{1}-{e}_{i\mathrm{a}}\frac{\mathrm{1}+R}{\mathrm{1}-R}\right)}^{m}.\end{array}$

The relation is plotted for different stress exposure amplitudes in Fig. 9.

Figure 9Impact of stress ratio on the relation between allowable cycles using any testing stress ratio and a stress ratio of $R=-\mathrm{1}$ (Eq. 6) shown for different stress exposure amplitudes. An S–N curve exponent of m=10 was used.

The pure testing time in days can be determined using

$\begin{array}{}\text{(7)}& {T}_{\mathrm{test}}=\frac{{N}_{\mathrm{test}}}{{f}_{\mathrm{test}}}\cdot \frac{\mathrm{1}\phantom{\rule{0.25em}{0ex}}\mathrm{day}}{\mathrm{86}\phantom{\rule{0.125em}{0ex}}\mathrm{400}\phantom{\rule{0.125em}{0ex}}\mathrm{s}},\end{array}$

where Ntest corresponds to the number of test cycles and ftest to the test frequency.

For a lead–lag FFST of the DTU10MW blade in a pitching position of $\mathit{\theta }=\mathrm{0}{}^{\circ }$ and a test frequency equal to the blade's eigenfrequency of ftest=0.965 Hz, a stress ratio of $R=-\mathrm{0.9}$ with an stress exposure amplitude of ea=0.25 and ${N}_{\mathrm{test}}=\mathrm{3.0}×{\mathrm{10}}^{\mathrm{6}}$ results in a testing time of Ttest≈36 days. If the blade is pitched at $\mathit{\theta }=-\mathrm{20}\mathrm{\dots }-\mathrm{40}{}^{\circ }$, the stress ratio is increased from $R=-\mathrm{0.9}$ to $R=-\mathrm{0.5}$ and gets closer to stress ratios from field load calculations (Fig. 8). To overcome too high tension strains (Fig. 7), the stress exposure amplitude is decreased to ea=0.125. Owing to the effect of a higher stress ratio the testing time can be reduced by 23 % to Ttest≈28 days (green dot vs. red dot in Fig. 9).

When performing an SCT of a blade segment at 10 % blade length, the stress ratio of $R=-\mathrm{0.9}$ is close to the field calculations (Fig. 8). For this case, the testing time cannot be reduced by the testing cycles Ntest but only by an increase in testing frequency. Assuming a frequency of ftest=1.0…1.5 Hz, the testing time is slightly to moderately reduced to Ttest≈35…23 days compared to an FFST at $\mathit{\theta }=\mathrm{0}{}^{\circ }$. When a segment at 70 % blade length is considered for an SCT with a realistic stress ratio of $R=-\mathrm{0.25}$ from field calculations (Fig. 8) and an stress exposure amplitude of ea=0.25, the test cycles can be further reduced by 47…68 % compared to an FFST with a pitch angle between $\mathit{\theta }=-\mathrm{20}\mathrm{\dots }-\mathrm{40}{}^{\circ }$ (green square vs. red dot in Fig. 9). Assuming a test frequency between ftest=1.0…1.5 Hz, the testing time results in Ttest≈11…7 days.

4 Conclusions

The simultaneous benefit and drawback of the SCT method is that only one blade segment is considered. Furthermore, the design of the boundary conditions of an SCT relies on detailed models.

There are different application scenarios for SCT, e.g., the detailed exploration of different design variants of a certain blade part under specific loading (not possible in FST). These experiments can be conducted within the design phase before or in hand with the certification FST or, in the case of a design revision, after the certification FST. Although these benefits of SCT can add value to the design process, SCT is not necessarily intended to replace FST.

On the basis of the assumptions presented for the damage calculation, it is shown that the mean stress level plays a substantial role in the fatigue damage calculation.

Data availability
Data availability.

The used input data and obtained result data are available at https://doi.org/10.5281/zenodo.1208135 (Rosemeier et al., 2018). The script necessary to reproduce the results is not publicly available since it partly refers to not published source code.

Appendix A: Derivation of Eq. (6)

After several rearrangement steps, Eq. (5) can be written as

$\begin{array}{}\text{(A1)}& {e}_{i\mathrm{m}}={e}_{i\mathrm{a}}\frac{\mathrm{1}+R}{\mathrm{1}-R}.\end{array}$

Furthermore, the initial steps deriving Eq. (6) are

$\begin{array}{}\text{(A2)}& \frac{{N}_{iR}}{{N}_{iR=-\mathrm{1.0}}}=\frac{{\left(\frac{\mathrm{1}-{e}_{i\mathrm{m}}}{{e}_{i\mathrm{a}}}\right)}^{m}}{{\left(\frac{\mathrm{1}}{{e}_{i\mathrm{a}}}\right)}^{m}}={\left(\mathrm{1}-{e}_{i\mathrm{m}}\right)}^{m}.\end{array}$

Equations (A1) and (A2) can be used to derive Eq. (6).

Competing interests
Competing interests.

The authors declare that they have no conflict of interest.

Special issue statement
Special issue statement.

This article is part of the special issue “Wind Energy Science Conference 2017”. It is a result of the Wind Energy Science Conference 2017, Lyngby, Copenhagen, Denmark, 26–29 June 2017.

Acknowledgements
Acknowledgements.

We acknowledge the support of the European Commission's Seventh Framework Programme within the IRPWind project (609795) and the support within the Future Concept Fatigue Strength of Rotor Blades project granted by the German Federal Ministry for Economic Affairs and Energy (BMWi) (0325939) and the Senator for Health, Environment and Consumer Protection of the Free Hanseatic City of Bremen within the ERDF programme Bremen 2014-2020 (201/PF_IWES_Zukunftskonzept_Betriebsfestigkeit_Rotorblätter_
Phase I).

Edited by: Athanasios Kolios
Reviewed by: Peter Joosse and one anonymous referee

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The span is measured from the blade root.

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