Introduction and
motivation
The cost of energy (CoE) is
the key parameter that determines the success of an energy source. In recent
years, both industry and the wind energy scientific community have focused
their efforts on the reduction in the CoE, with the goal of increasing the
competitiveness of energy from wind with respect to other technologies. A
reduction in the CoE can be obtained by a variety of means, one of the most
significant effects coming from an increase in the annual energy production
(AEP). AEP can be increased by improving the aerodynamic efficiency of the
rotor and by harvesting a greater amount of energy with larger swept areas
and taller towers. Because of this, together with other scale benefits
typically associated with larger wind turbines, there is a very clearly
marked trend towards bigger machines. In the offshore case, where logistics
and transportation are very different from onshore, the tendency towards very
large wind turbines is even clearer, the optimum size not having been reached
yet.
To satisfy this growth trend, the simple upscaling of existing machines is
unfeasible. In fact, as cost is typically well correlated with mass and mass
with volume, a naive scaling would translate into an unacceptable cubic
growth of cost. Among other approaches, load alleviation techniques help
address this issue, increasing the efficiency of the aerostructural
configuration and limiting the cost growth rate of wind turbine components
.
The mitigation of loads can be obtained by
full-span/distributed and passive/active
solutions. Full-span solutions involve the response of the entire blade.
Individual pitch control (IPC) is a full-span active technique, which is
seeing an ever increasing acceptance by industry, while bend–twist coupling
(BTC) is an example of the full-span passive category .
Although often very effective, any full-span solution is inherently somewhat
limited in bandwidth, due to the inertia and nonlocal response of the blade.
Moreover, pitch-bearing fatigue and the usage of the pitch actuators limit
the application of full-span active techniques. Distributed solutions, on the
other hand, locally affect the flow using flaps, tabs or other devices. The
local nature of these solutions allows in principle for a higher bandwidth
both in space and in time, which could potentially result in an even higher
reduction in loads. This should however be traded with their higher
complexity, which might in turn affect CoE because of higher production and
maintenance costs and/or decreased availability.
Numerous distributed active solutions for horizontal axis wind turbines
(HAWTs) have been explored, often inspired by aeronautical applications. At
present, the most mature applications appear to be the ones based on trailing
edge flaps
,
although alternative solutions based on micro-tabs and compliant
structures have also been considered .
Passive distributed techniques were first developed for aeronautical
applications. An early example of passive load alleviation is reported by
, in which a long-period overbalanced flap was used to
reduce airplane accelerations due to atmospheric gusts. A comparison of
different passive devices for the alleviation of vibratory loads on
helicopter rotors is described by . This study identified
the passive blade tip concept as the most promising technique to improve the
aeromechanical qualities of the rotor. Blades were modified in their
outermost portion to install a free pitching tip. The relative rotation
between tip and the inner blade was driven by the aeroelastic loads, and the
device parameters were tuned to achieve the desired dynamic response.
The design of the passive tip is the outcome of intense research activity at
NASA in the 80s. The simulation of a passive tip concept is described by
, aiming at a more uniform airload distribution during the blade revolution by self-adjusting
blade tips. Analytical results showed an improvement in lift generated by the
rotor in cruise conditions and a reduction in drag and the power required.
Since the mean relative rotation of the tip is related to the restraining
moment at the hinge, a preload was used as a tuning parameter to modulate the
blade tip angle of attack and the resulting aerodynamic forces. The passive
tip concept was also validated through experiments , which
confirmed a considerable reduction in required power in high thrust
conditions. This result is related to a favorable influence of the blade tip
negative pitch angle with respect to the inboard blade portion. Furthermore,
the flapwise and control loads were reduced considerably, although no
positive effects were observed on the lead–lag loads. Additional studies
focused on the configuration of a passive torque controller used to adjust
the preload . This fully passive mechanism converts
centrifugal loads in a preset torque at the movable tip. Including
considerations on simplicity and reliability, the most promising solution
appeared to be one that generates the output torque from the tensile loading
of two twisted wire straps .
Notwithstanding these promising results, passive tips have not been adopted
by the helicopter industry. Although active flaps have also not yet arrived
on the market, they have seen some significant demonstration by industry
. In fact, in aeronautical applications higher levels
of complexity are acceptable if they entail superior performance, weight
savings or the improvement of relevant metrics. Therefore, in this case
active flaps might be in general more interesting than passive ones. The
situation is different in the wind energy case, where the main (often unique)
driver is the CoE. In this case, availability and maintenance costs are of
paramount importance. From this point of view, deploying a wind turbine with
active flaps in the field still seems to be a very significant challenge.
Therefore, for wind energy applications a passive solution might be more
appealing than an active one if the former implies greater simplicity,
robustness and ease of repair than the latter.
Among the first applications of passive distributed solutions in wind energy
is the airfoil camber regulation described by . This
passive device adapts blade camber to alleviate pressure fluctuations. The
desired behavior is obtained by tuning the structural properties of the
device, consisting of a spring and damper. A preliminary validation was
performed on a typical section model, while a more recent analysis is
reported in , where a nonlinear lifting-line free-vortex
wake model is employed to assess the performance of the passive device on a
multi-MW HAWT. Results indicate a reduction in the standard deviation of
blade root bending moments, although a single simulation was considered.
Another passive camber solution is based on bistable composite structures
. In this case, the airfoil camber variation is triggered
by the aerodynamic loads that modify the equilibrium condition of a compliant
structure with embedded multi-stable elements. This technique results in a
discrete control action because only a finite number of stable
configurations are possible. Furthermore, an external load has to be provided
to restore the original blade camber.
A fully articulated passive flap was first proposed by .
The idea is in this case to offset the flap center of gravity forward of the
hinge line. This way, flapwise accelerations of the blade excite a response
of the flap that, by changing the airfoil camber, tends to oppose the
acceleration itself, thereby attenuating blade loading and in turn fatigue.
The flap is also aerodynamically balanced, in the sense that it is designed
not to respond to the deliberate changes in the angle of attack imposed by
the wind turbine control system. Multiple load cases were considered through
a loose coupling procedure based on a state-of-the-art aeroservoelastic
simulator and a typical section model, indicating very promising performance.
As the literature shows, a few recent studies have considered passive flaps
for HAWTs. However, one of the most promising solutions for rotorcraft
applications, the blade free tip, was only explored and tested during the 90s
by the FLEXHAT program . That study considered
a two-bladed HAWT equipped with an elastomeric teeter, a flexbeam allowing a
limited flapping motion, as well as two passively activated blade tips.
Despite the considerable advantages in terms of load reduction, the solution
was not developed further due to its inherent complexity.
The present study tries to further explore the tip concept, investigating
various solutions for the alleviation of loads on multi-MW HAWTs (see
Fig. ). Passive, semi-passive
and active solutions are considered in order to provide a general overview of
the possible range of configurations and their respective performance. The
passive solution is purely activated by aerodynamic loads, while the
semi-passive one uses an active component to apply a varying restraining
torque to limit mean tip deflections according to the machine operating
condition. Finally, the active solution uses an actuator to drive the tip
deflection based on a feedback control law. Each configuration is analyzed in
detail, including the tuning of the respective parameters. Performance is
assessed using the accepted international certification standards within a
high-fidelity aeroservoelastic simulation environment.
Articulated blade tip concept for load
alleviation.
![]()
The paper is organized as follows. Section
considers the tip design problem. Passive and semi-passive configurations are
examined first, providing some general guidelines and a preliminary sizing of
the main system parameters for the aeroelastic integration of the devices on
board the wind turbine. The active solution is then introduced, and its
control algorithm is tuned. Next, Sect.
compares fatigue and ultimate loads as well as off-design conditions.
Finally, conclusions and an outlook on possible future developments are
reported in Sect. .
Design of blade tip devices
The design of the blade tip focuses here on the properties of the hinge
connecting it to the rest of the blade, while the external blade shape is
kept constant. This simplification distinguishes the effects of the tip
motion per se from further possible effects that could be obtained by
modifying its aerodynamic shape. While the approach might be suboptimal, a
specific tailoring of the aerodynamic characteristics of the tip can be
analyzed at a later stage.
Passive and semi-passive configurations
The device design aims at optimizing the tip motion in order to mitigate
loads. The positions of the hinge line (HL), of the tip aerodynamic center
(AC) and of the center of gravity (CG) (see
Fig. ) play a crucial role in
determining the physical phenomena contributing to load alleviation.
Wind turbine blade with articulated tip.
![]()
If the hinge line is close to the aerodynamic center of the blade tip, then
the aerodynamic moment is nearly independent of angle of attack changes.
Therefore, the device behavior is mainly driven by the inertial response of
the blade tip if its center of gravity is offset with respect to the hinge.
This is the same load alleviating mechanism used by for
their passive trailing edge flap. By contrast, if the hinge line is away from
the blade tip aerodynamic center while the center of gravity is not, then the
response is mainly driven by aerodynamic loads. In particular, when the hinge
line is forward of the aerodynamic center, an increase in the angle of attack
at the blade tip will induce an increase in lift and, consequently, a nose
down moment at the hinge that will induce a pitch down rotation. This will
eventually oppose the original increase in the angle of attack, thereby
realizing a load-mitigating action.
Both the inertial and aerodynamically driven solutions can be used for
designing passive load-mitigating devices. However, while the former proved
to be very effective for the flap case , the latter
seems to be better suited to the tip case considered in the present study.
Several factors make the inertial-driven solution difficult to implement for
a tip device. First, a flap is characterized by the hinge moment rate of
change with respect to both angle of attack and flap deflection changes, two
parameters that can to a large degree be set independently of each other. By
contrast, a tip device is only characterized by its sole hinge moment rate of
change with respect to the angle of attack; in addition, the moment with
respect to the aerodynamic center is not null because of the non-null camber
of the tip airfoils. Therefore, it is much harder for the tip case to obtain
a good alleviating performance and small sensitivity to disturbances such as
gravity and centrifugal loading. In addition, a significant mass ballast is
needed to obtain the necessary inertial effects, ballast that in turn lowers
the blade natural frequencies and may negatively affect loading. Based on
these considerations, the aerodynamically driven solution is adopted for the
present study.
The hinge location is a compromise between the weathercock tendency of the
blade tip, which suggests a forward position, and a desire to limit inertial
couplings, which suggests a hinge position close to the center of gravity of
the tip.
The spanwise extent of the blade tip was optimized with the help of a
parametric analysis, considering a trade-off among blade root load
alleviation, loading at the hinge and impact on power capture.
The wind turbine is operated with a variable-speed pitch–torque control
strategy, including the partial load regime (or region II) from cut-in to
rated speed and the full-load regime (or region III) from rated to cut-out
speed . The best possible aerodynamic performance is
sought in region II to optimize power capture. Therefore, the mean
misalignment of the tip with respect to the rest of the blade should be as
small as possible not to negatively affect the rotor efficiency. On the other
hand, an excess of power is available in region III, so that a mean
misalignment of the tip is permissible in this case as it would be readily
compensated for by the control system without incurring any AEP loss.
A torsional spring and torque preload are used at the hinge with the aim of
controlling the tip response. The tip pitch dynamic equilibrium is
Jθθ¨+Kθθ=Mp+Ma,
where θ is the tip pitch rotation, Jθ the tip inertia,
Kθ the torsional spring stiffness at the hinge, Mp the
hinge preload and Ma the aerodynamic moment. The primary device
design parameters are Kθ and Mp.
The torsional spring Kθ was calibrated to limit the tip pitch
oscillation amplitude. This tuning was performed by running aeroservoelastic
simulations in steady and turbulent conditions for varying wind speeds
spanning the entire operating range of the machine and identifying an optimal
compromise between fatigue alleviation and power loss. Although in principle
the spring stiffness might be scheduled with respect to the operating
condition, it was found that a constant average value was a simpler and
similarly effective solution.
The tip mean misalignment is controlled by providing a torque preload
Mp at the hinge. As the aerodynamic loading at the tip, and hence
its mean moment at the hinge, depends on the operating condition, the preload
should be varied on account of the operating point at which the machine is
functioning. In the semi-passive configuration the preload is generated by an
actuator, while in the passive case by a mechanical device that produces a
torque in response to the centrifugal loads generated by the blade rotation.
In both cases, the resulting preload at the hinge is directly related to the
rotor angular speed.
A sketch of the passive and semi-passive configurations is found in
Fig. .
Top: semi-passive tip configuration. Bottom: passive tip
configuration.
![]()
As the preload is related to the operating point, its value can be computed
in steady-state normal wind profile (NWP) conditions using a complete
aeroservoelastic model of the wind turbine, scanning wind speeds from cut-in
to cut-out. To speed up the identification of the necessary preload value, at
each wind speed a simulation was run where the relative rotation in the tip
hinge was set to zero. Once the solution had settled into a periodic cycle,
the mean value of the resulting torque in the hinge was used as the preload
value for that operating condition.
In principle, the preload could be scheduled with respect to the mean wind
speed or to the rotor angular velocity. The former option is more complicated
and possibly less reliable because it requires an observer to estimate the
rotor-equivalent wind speed. By contrast, scheduling the preload with
respect to the rotor angular velocity is simpler, since measurements of the
rotor speed are available on board wind turbines. As the angular velocity is
constant in region III, a constant preload above rated wind speed will result
in a non-null mean misalignment of the tip. This is not a problem, as there
is a power excess in this condition, so that a less efficient rotor does not
pose any concern. The situation would be different for a machine with a
transition region II1/2 in between regions II and III – which happens
whenever the rotor speed hits its upper limit before rated power is reached
– where scheduling with respect to rotor speed alone might incur power
losses.
For the semi-passive configuration, an actuator applies the necessary preload
torque at the hinge based on a look-up table storing the load-rotor speed map
Mp=Mp(Ω) obtained in the previously described
analyses, where Ω is the rotor angular velocity. No feedback
regulation is involved, and the actuator simply uses the filtered (to remove
fast fluctuations and noise) rotor speed as feedforward information.
The passive configuration uses centrifugal forces caused by the rotor angular
rotation to generate the necessary preload Mp, without using
active components. To this end, in this paper we consider the mechanical
device described in , characterized by a screw joint that
relates any linear displacement z of the tip parallel to its hinge axis to
a corresponding rotation θ about the same axis, i.e., z=τθ,
where τ is the screw joint helical pitch or transmission ratio. The
actual mechanical design of this device is beyond the scope of this study,
and its characterization is here limited to the evaluation of its parameter
τ. The passive tip pitch dynamic equilibrium can be written as
JPTθ¨+KPTθ=τFc+Fg+Ma,
with
JPT=Jθ+τ2m,KPT=KPTθ+τ2Kz,
where JPT and KPT are the total inertia and torsional
stiffness of the passive tip device. These include the proper inertia of the
tip Jθ and the hinge spring KPTθ, in addition to
terms contributed by the screw joint, m being the tip mass and Kz the
screw linear displacement stiffness. In Eq. (), Fc and
Fg are the centrifugal and gravitational forces, respectively,
projected onto the blade spanwise direction. To a first approximation, the
effects due to blade out-of-plane and in-plane motion, as well as the
contributions of rotor cone and uptilt angles, are neglected. Therefore, the
centrifugal force is expressed as
Fc=mr+zΩ2,
where r is the radial position of the tip center of gravity. Inserting
Eq. () into Eq. (), one gets
JPTθ¨+KPT-mτ2Ω2θ=τmrΩ2+τFg+Ma.
The gravitational force is
Fg=mgcosψ,
where ψ is the blade azimuthal position and g the acceleration of
gravity. Since Fg is a periodic disturbance with zero mean over a
revolution, the transmission ratio τ is chosen such that the first term
on the right-hand side of the equation balances the aerodynamic moment at the
hinge line, leading to
τ=-MamrΩ2.
An average value of Ma over the most likely operating conditions
(between 7 and 9 ms-1, according to the Weibull distribution considered here) is used to compute τ.
The value of the hinge spring stiffness for the passive tip case was set by
requiring this device to have the same modal frequency as the semi-passive
case, which is readily computed from Eq. () as ωSP2=Kθ/Jθ. By setting Kz=0 and using Eq. (), one
gets
KPT,θ=ωSP2JPT+mτ2Ω2≈ωSP2JPT,
where the term depending on angular velocity was dropped because it was negligible.
This choice results in a hinge stiffness that, conveniently, does not depend
on the operating condition, as in the semi-passive case.
It should be stressed that this is not the only possible criterion to
determine the hinge spring stiffness for the passive tip case. In fact, the
tip mode could in principle be placed anywhere in the spectrum, as long as it
does not create resonant conditions with the per-rev harmonic excitations and
with other natural frequencies of the machine. On the other hand, the present
approach seemed to work well in practice. In fact, raising this frequency by
increasing the spring stiffness limits the tip pitch oscillations, in turn
reducing its authority. The opposite approach of lowering the frequency by
softening the spring has the effect of increasing the disturbance caused by
gravity. In fact, gravity cyclically pulls on the blade tip, creating a
radial displacement that, through the screw joint, induces a pitch rotation,
which in turn creates a 1P disturbance. The present approach was found to
provide a good compromise between these two contrasting requirements,
although a further fine-tuning of the parameters is probably still possible.
Reference wind turbine and simulation environment
The blade tip devices are sized and studied with application to the 10 MW
Reference Wind Turbine (RWT), developed by Danmarks Tekniske Universitet
(DTU) . Some of the principal parameters of the machine
are reported in Table , while the full database
can be downloaded at the project website .
Principal parameters of the DTU 10 MW
RWT.
Parameter
Value
Rated power
10 MW
Wind class
IEC 1A
Rotor diameter
178.3 m
Hub height
119.0 m
Rated wind speed
11.4 ms-1
All simulations are performed with an aeroservoelastic model of the wind
turbine implemented with the flexible multibody program Cp-Lambda
(see , and references therein). The baseline regulation
strategy is provided by an external library implementing the control routines
reported in . Based on a parametric study, the spanwise tip
length was set to 15 % of the blade length, while the tip hinge line was
located at 19.7 % of the local blade chord from the leading edge. The tip
is connected by a revolute joint to the rest of the blade for the
semi-passive and active configurations. In both cases, the hinge rotation is
driven by an actuator, modeled as a second-order system. For the passive
case, the tip is connected to the blade by a screw joint. In all cases, tip
excursions are limited to ±20∘ by unilateral contact conditions
in the joint.
It was decided to operate all devices both in the partial- and full-load
regions. Although operation in partial load conditions may affect power
output, there is also a significant accumulation of fatigue damage around
rated wind speed for sites with low most-probable winds. However, it is also
clear that, depending on the characteristics of machine and site, the sole
operation in full-load conditions might be preferable.
The aerodynamic model is based on standard blade element momentum (BEM)
theory , which is customarily used for the analysis of
loads on wind turbines. This approach is not able to model the shed vortex at
the blade-tip junction . This vortex induces a velocity
field that modifies the orientation of the lift vector at the tip, in turn
causing a variation of induced drag . Since in a passive
solution aerodynamic loads contribute to the motion of the tip device, this
modeling deficiency may also affect the accuracy of the predicted pitch
motion. This is one of the main approximations of the present work, which
will require further investigation by more sophisticated aerodynamic models.
Sizing of the passive and semi-active solutions
The wind turbine operating range is first analyzed in NWP conditions
. The associated rotor speed and blade pitch settings vs.
hub-height wind speed are shown in
Fig. .
Rotor speed Ω and blade pitch β vs. hub-height wind
speed V.
![]()
Nominal values of the torque preload Mp as a function of wind
speed were obtained by constraining to zero the tip rotation at the hinge and
measuring the resulting internal moment. The result is shown at the top of
Fig. , using a dash–dotted
line: by prescribing this preload at the hinge, one would obtain a zero mean
misalignment of the blade tip. Since the preload is adjusted based on rotor
speed both for the passive and the semi-passive configurations, this
reference preload can be followed only between 7 ms-1 and rated
wind speed, when the rotor speed does indeed change (see
Fig. ).
Top: preload Mp at the hinge line vs. hub-height wind
speed. Bottom: hinge stiffness Kθ vs. hub-height wind speed.
![]()
As shown in the figure, for lower and higher wind speeds the actual preload
provided by the passive and semi-passive solutions remains constant, implying
that the blade tip will have a non-zero mean pitch offset with respect to the
blade. The preload can be actively changed by a torque actuator in the
semi-passive tip solution, so that the provided preload exactly follows the
nominal one in this case. For the passive configuration, the preload is
obtained by a constant transmission ratio τ connecting tip spanwise
displacements with tip pitch rotations, which, as shown in the figure, still
approximates very well the nominal preload behavior.
Figure on the bottom shows with a dash-dotted line the hinge spring stiffness that would result in a ±10∘
oscillation of the tip in NWP conditions. As this stiffness changes little
with respect to wind speed, it was approximated with a constant value for the
semi-passive case, further tuned with the help of turbulent analyses. From a
practical point of view, a constant spring stiffness is useful because it
reduces the complexity of the device. As previously explained, the hinge
stiffness for the passive configuration differs from the one of the
semi-passive case. In fact, since the transmission ratio of the screw joint
increases the torsional inertia of the blade tip, the hinge stiffness was
increased to keep the tip mode at the same frequency in both solutions.
Table reports the modal
frequencies of the rotating blade in a vacuum at rated speed, for the
baseline blade and the semi-passive and passive solutions. Minor differences
are due to the adoption of a constant transmission ratio τ, which
however is important for the simplicity of the device. The blade tip mode is
clearly distinct from the lower blade frequencies, limiting the risk of
aeroelastic interactions.
Modal frequencies of the rotating blade in a vacuum (in
rads-1).
Mode
Baseline
Semi-passive tip
Passive tip
First flap
4.08
3.99
3.99
First edge
5.67
5.40
5.39
Second flap
10.3
10.9
10.8
Second edge
15.6
16.0
15.9
Third flap
20.0
21.9
21.9
Tip mode
–
25.2
25.1
Third edge
31.2
33.0
32.6
Active configuration
Besides the passive solutions described earlier, tips can also be used for
active feedback control. In this case, pitch motions are actively driven by
tip actuators. Due to the lower inertia of the tip with respect to the entire
blade, tip-based active control might have a higher bandwidth than full-span
pitch control. In addition, as the tip has a high moment arm with respect to
the blade root, even relatively small changes in the aerodynamic loads might
have significant repercussions on the overall loading of the rotor. Both of
these effects might be especially visible for larger turbines, although a
detailed investigation of scale effects is beyond the scope of the present
work.
In this paper, cyclic pitch control of the tips is used for the reduction of
rotor moments in the fixed system,
using a formulation similar to one used for classical full-span IPC
.
Blade bending moments are measured by load sensors at the blade roots and
transformed first into out-of-the-rotor-plane moments and then into direct
Md and quadrature Mq moments in the fixed frame by
the Coleman transformation .
After filtering to remove frequencies at and above 1P, reference loads are
subtracted from the Coleman-transformed moments, yielding the delta loads
used for feedback ΔM=M-M*(V‾) for both the q and r
components, where V‾ is a slowly varying moving average of the
wind speed used for scheduling the reference loads. The use of delta loads is
useful because of the lower authority of tip pitch control compared to
full-span pitch control. In fact, by cyclically pitching the whole blade,
full-span pitch control can very significantly reduce the mean value of
fixed-frame loads, which is typically not possible with the sole use of tips.
A proportional–integral (PI) controller is then formulated in the fixed
frame, giving
β=kPΔM+kI∫0tΔMdt,
where kP and kI are the proportional and integral
gains, respectively. The same control law is used for the direct and
quadrature components, yielding both the βd and
βq control inputs in the fixed frame, which are finally
transformed back into the rotating system via Coleman's inverse transform.
This IPC formulation results in a 1P tip pitch input activity.
Higher-frequency Coleman transformations could be easily used within the
exact same technique to obtain a higher-harmonic
controller. In fact, given the reduced inertia of an active tip device, a
wider bandwidth control activity could be more easily achieved than using
full-span pitch control, especially for very heavy and large blades.
However, a fatigue analysis performed on the reference wind turbine
considered in the present study revealed that fatigue is primarily generated
in a very low range of frequencies. In fact, Fig. reports the
normalized blade root lifetime bending moment damage-equivalent load (DEL) as
a function of load harmonics for the baseline RWT. It appears that DEL
increases very rapidly with frequency, to the point that 75 % of
damage is already accumulated for frequencies up to 1P. Damage then rapidly levels
off, with very little contributions coming from frequencies above the 3P. For
this reason, and given the preliminary nature of the present study, it was
decided to limit the tip control activity here to the sole 1P harmonic.
Normalized blade root lifetime DEL bending moment, plotted as a
function of load frequency.
![]()
Tuning of the active tip control law
Tuning of the cyclic tip pitch controller involves setting the reference
values for the direct and quadrature loads, as well as the proportional and
integral gains.
Figure shows the Md (at top)
and Mq (at bottom) values vs. wind speed for the baseline wind
turbine without tips, using dash–dotted lines. The same figure also shows the
reference values Md* and Mq*, using dashed
lines. These values were chosen by trial and error and, as previously
explained, aim at lowering the feedback loads due to the reduced authority of
a tip compared to a full-span pitch control solution.
Top: average Md load (dashed line) and reference value
Md* (dash–dotted line) vs. wind speed. Bottom: average
Mq load (dashed line) and reference value Mq* (dash–dotted line) vs. wind
speed.
![]()
The tip controller was tuned using turbulent wind conditions (DLC (design load case) 1.1,
). Gains were set in order to achieve satisfactory
performance on hub loads, while at the same time avoiding an excessive
actuator duty cycle (ADC). The performance of the controller was checked with
respect to the most demanding conditions (DLC 1.3). A simple gain scheduling
was used to further boost performance, by multiplying the gains by a factor
of 4 around rated wind speed, and specifically between 9 and 11 ms-1.
Due to the lower loads sustained by tip actuators compared to blade root
ones, tip IPC was used over the whole operating range of the machine and not
only in region III as customarily done for full-span blade IPC.
Both the reference loads and gains were scheduled using a 30 s
moving-averaged wind speed measured from the nacelle anemometer. Fixed-frame
loads were low-pass filtered with a fourth-order Butterworth filter with a
cut-off frequency of 0.1 Hz.
Results
The performance of the proposed tip devices was evaluated by studying the
wind turbine in different operating conditions, as recommended by
international certification standards . Of all various DLCs
used to design the machine , the most demanding ones in
terms of fatigue and ultimate loads were selected. In turbulent wind
conditions, results were averaged over four different realizations
corresponding to different seeds .
Standard design conditions
The standard power production range was simulated by DLC 1.1 from the cut-in
to the cut-out speeds in 2 ms-1 increments, without the
extrapolation of loads.
The AEP percent variations with respect to the baseline configuration without
tip devices are reported in Fig. . Apparently,
the active tip device has the largest impact on energy capture, possibly due
to the choice of operating it also in region II. The maximum AEP reduction is
equal to 0.5 %, a value that is not insignificant and could offset the
advantages in terms of load alleviation. A complete redesign study should be
used to combine the variation of AEP with load reductions into a single CoE
value. Note, however, that for the passive and semi-passive solutions AEP
reductions are very small and possibly within the margin of accuracy between
any two of these simulation analyses.
Percent AEP variation with respect to baseline configuration.
![]()
DELs were evaluated at a number of points on the machine based on rainflow counting. The blade, main
bearing and tower base were selected as fatigue verification spots because
they are indicative of possible structural regions prone to fatigue problems.
DELs corresponding to the combined moment at the most damaged point at each
verification section are reported in
Fig. .
Top: percent variation of DELs at verification spots. Bottom:
percent variation of DELs vs. blade span.
![]()
The effects of the appended devices at the blade root, main bearing and tower
base are shown in the top part of the figure. All three tip devices appear to
be lowering fatigue loads, although to a different extent at different
verification spots. The active tip achieves the best load reduction at the
main bearing because those are indeed the loads targeted by the tip IPC
control algorithm. On the other hand, it is interesting to observe that the
passive and semi-passive tips perform better than the active configuration at
the tower base, where the DEL is mainly due to rotor thrust. In fact, these
results seem to indicate the ability of the passive and semi-passive tips to
smooth out load fluctuations due to turbulence. As the three tips operate
independently, in contrast with the centralized operation of the IPC
algorithm, they are better able to react locally to local wind fluctuations,
in turn resulting in smaller fatigue damage at the tower base. The effects at
the blade root are also significant, the semi-passive achieving the best results,
followed by the active tip and finally closely followed by the fully passive
configuration.
However, a more detailed analysis of blade fatigue reveals significant
differences among the three solutions, as shown in the bottom part of
Fig. . In particular, the plot of
DEL vs. blade span shows that the passive and semi-passive solutions reduce
fatigue throughout the whole span of the blade, which again indicates the
ability of the tips to smooth out aerodynamic loads. By contrast, the
active tip lowers fatigue towards the root, but increases it at the tip. This
is due to the commanded pitch activity that, with the final goal of lowering
nodding and yawing moments at the main bearing, in reality overloads the
blade tip. Usually fatigue may become a design driver in the inner portion of
the blade, so the increase in DEL towards the tip might not be a major source
of concern. Nevertheless, a rise in fatigue damage in the tip region should
be expected during blade design and would have to be considered.
Pitch activity is reported in Fig.
in terms of ADC vs. hub-height wind speed, where ADC over a time span T is
computed as
ADC=1T∫0Tβ˙dt.
The collective pitch ADC, which measures the blade pitch activity performed
by the controller governing the machine , is shown in the
top part of the picture. Differences are modest, with some reduction
noticeable for the semi-passive solution. This can be attributed once again
to the smoothing of the airloads
performed by this device, which in turn yields a smoother response of the
machine and a consequent slightly reduced activity of the controller in
reaction to wind fluctuations.
Top: blade pitch ADC vs. wind speed. Bottom: tip ADC vs. wind
speed.
![]()
The bottom part of Fig. shows the
tip ADC vs. wind speed. For the passive and semi-passive solutions, ADC is
only a measure of how much the tip pitches in response to load fluctuations,
while for the active case it represents a measure of the control effort
exerted by the actuator. The plot shows that the three devices have very
roughly similar tip activities, although these are in nature quite different,
as shown by the previous load analysis. In addition, it appears that the
semi-passive device has a more pronounced activity than the passive one.
An ultimate load analysis was performed by considering a selected set of
DLCs. DLC 1.1 and 1.3 consider power production in standard and extreme
turbulence conditions. In DLC 2.3, a deterministic gust occurs in conjunction
with a grid loss, and the effects of the fault time are examined by multiple
simulations. Finally, DLC 6.2 considers parked conditions with grid failure,
where multiple yaw conditions are considered to identify the worst scenario.
Attention is focused on the combined bending moments at the blade root, main
bearing and tower base, and percent variations of the ultimate loads with
respect to the baseline are reported in
Fig. . Better performance
is achieved at the main bearing and at the blade root, where the most
demanding situations are due to DLC 1.3. Here again, as in the case of
fatigue damage, the tip devices seem to be able to smooth out airloads, with a beneficial effects
also on peak loads.
Percent variation of ultimate loads at verification spots with
respect to baseline.
![]()
The situation is different for ultimate loads at the tower base, which are
due to DLC 6.2. Although in this case tip oscillations do not in general help
in reducing loads, the ability of the active and semi-passive solutions to
deflect the tip can be used to gain a modest advantage. In fact, by pitching
the tip one may reduce the sail area of the blade, which in turn may somewhat
reduce loads during storms. For these two cases, tips were pitched all the
way to their stop positions (20∘). As shown by the figure, this
strategy results in a modest decrease of loads at the tower base. This active
protection of the rotor in storm conditions is not possible with the fully
passive solution, where the tip is free to float into the wind but cannot be
controlled directly. The same figure shows that this has a very modest
negative effect on tower loads.
Off-design conditions
The effects of a blade tip failure are investigated to understand if the
advantages of the proposed tip devices can be offset by a fault of the tip
pitching system. The fault is investigated by blocking the relative rotation
of a single blade tip, while the other ones are functioning in a regular way.
It is supposed that the wind turbine is equipped with a safety system to
detect the fault and trigger an immediate shut down procedure. Generator
fault or loss of electrical network are not included in the fault scenario because simultaneous malfunctions are considered as very unlikely.
Blade tip faults are examined using DLC 2.1 and 2.3 to identify the most
critical condition. A single seed is considered for DLC 2.1 normal
turbulence model (NTM) simulations
because the relative position of the fault with respect to wind fluctuations
is more important than the analysis of different wind realizations. The blade
tip fault is imposed in conjunction with a positive steep gradient or a
maximum of the hub-height wind speed. These two conditions are respectively
labeled “grad” and “peak” in the following. When turbulent winds are
considered, each simulation is associated with a number, which represents the
mean hub-height wind speed, and a letter, identifying a turbulent seed.
DLC 2.3 simulates a deterministic extreme operating gust (EOG) at cut-out
(labeled vo), rated (labeled vr) and rated ±2 ms-1 (labeled
vr ± 2) wind speed. In total, 16 simulations were performed at each
wind speed, varying the time interval between the gust and the fault as well
as the azimuthal position of the faulty blade tip. Each simulation is
identified by a number that refers to one of these combinations.
The off-design performance is investigated by ranking the ultimate loads of
the standard envelope plus the fault conditions in decreasing order and
monitoring the variation of the maximum load magnitude. The first three
ranking combined moments at the main bearing are reported for each
configuration in Fig. , where the
fault conditions are identified by using gray-shaded bars.
Ranking analysis of main bearing combined moment. Blade tip fault
conditions are displayed using gray-shaded
bars.
![]()
A blade tip failure is considered dangerous if the maximum load magnitude
increases with respect to the baseline configuration. The ranking analysis
for blades and tower base are not reported here because fault conditions do
not modify the five highest-ranking loads. In fact, DLC 1.3 remains the load
case driving blade design, while the tower is still stressed by DLC 6.2. By
contrast, the combined moment at the main bearing is affected by the rotor
imbalance caused by the blade tip fault. Therefore, off-design conditions may
generate loads that are comparable to, or even higher than, in the non-faulty
standard DLCs.
The results reported in the figure show that all tip devices do not exceed
the load envelope of the baseline machine. In addition, fault conditions are
not load drivers for the passive and semi-passive solutions, while they
produce the leading load for the active tip case. This might be due to the
loss of coordination of the blade tip movement that follows a tip fault.
Conclusions and future work
A movable blade tip concept for load mitigation on wind turbines has been
investigated in this paper. Although already studied for rotorcraft
applications, the installation on multi-MW HAWTs is, to the authors'
knowledge, a novelty. The device allows for a relative pitching motion of the
blade tip with respect to the rest of the blade, introducing a further
control capability.
Passive, semi-passive and active blade tip solutions were developed and
compared. The passive solution achieves the simplest configuration because it
does not involve any active component. The active tip requires sensors and
servomotors and implements a feedback control algorithm. The semi-passive tip
is in a sense in between the other two configurations, requiring active slow
regulation of the hinge preload but no feedback control. The free motion of
the passive and semi-passive devices is driven by the weathercock tendency of
the tip due to a suitable chordwise location of its hinge, together with a
restraining spring. These devices result in a passive decentralized control
strategy powered by local fluctuation of the aerodynamic loads. The resulting
tip pitching smooths out the airloads without incurring in AEP losses of any significance. By contrast,
the active tip implements a centralized IPC control strategy, that targets
the nodding and yawing moments at the hub and seems associated with a more
significant impact on AEP.
The paper described the preliminary sizing of all devices. The hinge preload
and stiffness for the passive and semi-passive configurations were defined by
an ad hoc procedure, while simple guidelines were reported for the tuning of
the gains of the active tip system.
The devices were tested in a comprehensive simulation environment, with
application to a large conceptual future machine. The analysis considered
both fatigue and ultimate loads, including also tip fault conditions,
following accepted standard certification guidelines.
Based on the results of the present analysis, the following conclusions may
be drawn:
All proposed tip devices improve on the baseline both in terms of fatigue
and ultimate load alleviation, although to a different extent on different
wind turbine components. These results might possibly be further improved by
a more complete optimization of the devices, including their aerodynamic
shape.
The more significant effects on fatigue are reported at the blade root
and tower base. For the passive and semi-passive devices, this seems to be
attributable to a smoothing of the airloads. Ultimate loads see the largest decrease at the main bearing,
while they are essentially unaffected on blade and tower.
None of the devices seems to significantly interfere with the collective
pitch–torque control system used for regulating the machine, although no
retuning of the controller was performed. For the semi-passive solution, the
load smoothing generated by the tip results in a slightly reduced duty cycle
of the blade pitch actuator.
The consequences of a blade tip fault are limited, with no effect on
the ultimate design-driving loads. The active and semi-active devices can be
used to reduce blade sail area in storm conditions. Although this technique
did not reduce ultimate loads on this specific machine, it might be
beneficial to other wind turbines more significantly driven by storm
conditions.
Further studies are clearly necessary before final conclusions may be drawn,
although these initial results seem to be promising. In particular, the
passive and semi-passive solutions behave nearly as well as the active one,
at a reduced complexity. This might be interesting for applications where
reliability, low cost of maintenance and high availability are important,
as in the offshore case.
The blade tip concept could be further developed along different lines. The
detailed design of the tip joint should be performed, addressing some
critical aspects such as the realization of the passive screw joint or the
installation of the servomotors. More sophisticated aerodynamic models could
be used to take into account the mutual interference between the tip and the
inner part of the blade, as well as the vortices shed by the twist
discontinuity at the joint . The control system could be
retuned to better account for the presence of the tip devices, while
shutdown procedures could also be revisited at least in the active tip case.
Finally, the integration of the blade tip concept in a rotor redesign
activity could shed light on the actual potential
beneficial effects on CoE or the lack thereof.