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Research Papers

Assessing the Importance of Geometric Nonlinear Effects in the Prediction of Wind Turbine Blade Loads

[+] Author and Article Information
D. I. Manolas

School of Mechanical Engineering,
National Technical University of Athens,
9 Heroon Polytechneiou Street,
GR15780 Athens, Greece
e-mail: manolasd@fluid.mech.ntua.gr

V. A. Riziotis

School of Mechanical Engineering,
National Technical University of Athens,
9 Heroon Polytechneiou Street,
GR15780 Athens, Greece
e-mail: vasilis@fluid.mech.ntua.gr

S. G. Voutsinas

School of Mechanical Engineering,
National Technical University of Athens,
9 Heroon Polytechneiou Street,
GR15780 Athens, Greece
e-mail: spyros@fluid.mech.ntua.gr

Contributed by the Design Engineering Division of ASME for publication in the JOURNAL OF COMPUTATIONAL AND NONLINEAR DYNAMICS. Manuscript received February 4, 2014; final manuscript received May 13, 2014; published online April 2, 2015. Assoc. Editor: Carlo L. Bottasso.

J. Comput. Nonlinear Dynam 10(4), 041008 (Jul 01, 2015) (15 pages) Paper No: CND-14-1033; doi: 10.1115/1.4027684 History: Received February 04, 2014; Revised May 13, 2014; Online April 02, 2015

As the size of commercial wind turbines increases, new blade designs become more flexible in order to comply with the requirement for reduced weights. In normal operation conditions, flexible blades undergo large bending deflections, which exceed 10% of their radius, while significant torsion angles toward the tip of the blade are obtained, which potentially affect performance and stability. In the present paper, the effects on the loads of a wind turbine from structural nonlinearities induced by large deflections of the blades are assessed, based on simulations carried out for the NREL 5 MW wind turbine. Two nonlinear beam models, a second order (2nd order) model and a multibody model that both account for geometric nonlinear structural effects, are compared to a first order beam (1st order) model. Deflections and loads produced by finite element method based aero-elastic simulations using these three models show that the bending–torsion coupling is the main nonlinear effect that drives differences on loads. The main effect on fatigue loads is the over 100% increase of the torsion moment, having obvious implications on the design of the pitch bearings. In addition, nonlinearity leads to a clear shift in the frequencies of the second edgewise modes.

Copyright © 2015 by ASME
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Figures

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Fig. 1

Co-ordinate systems definition

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Fig. 2

Dynamics of the infinitesimal beam element

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Fig. 3

Wind turbine inertial system and local components systems

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Fig. 4

Realization of multibody kinematics

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Fig. 5

Realization of multibody kinematics at the level of the component

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Fig. 6

Mean blade tip deflections

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Fig. 7

Blade torsion angle, for uniformly distributed load of 10 kN/m, acting in the flapwise direction and applied on the elastic axis

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Fig. 8

Blade extension, for uniformly distributed load of 10 kN/m, acting in the flapwise direction and applied on the elastic axis

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Fig. 9

Blade flapwise deflection, for uniformly distributed load of 10 kN/m, acting in the flapwise direction and applied on the elastic axis. 1st order model lies on top of 2nd order.

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Fig. 10

Bending–torsion coupling effect due to high bending deformation

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Fig. 11

Blade torsion angle versus flapwise deflection, for uniformly distributed load acting in the flapwise direction, applied on the elastic axis and ranging from 1 kN/m to 10 kN/m (each symbol on the lines corresponds to a step of 1 kN/m)

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Fig. 12

Blade torsion angle, for uniformly distributed load of 10 kN/m, acting in the flapwise direction and applied on the mass centre

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Fig. 13

Blade extension versus flapwise deflection, for uniformly distributed load acting in the flapwise direction, applied on the mass center and ranging from 1 kN/m to 10 kN/m (each symbol on the lines corresponds to a step of 1 kN/m)

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Fig. 14

Blade torsion angle versus flapwise deflection, for uniformly distributed load acting in the flapwise direction, applied on the mass center and ranging from 1 kN/m to 10 kN/m (each symbol on the lines corresponds to a step of 1 kN/m)

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Fig. 15

Blade torsion angle, for uniformly distributed combined flapwise and edgewise load of 10 kN/m (in each direction), applied on the elastic axis

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Fig. 16

Blade torsion angle versus flapwise deflection, for uniformly distributed combined flapwise and edgewise load, applied on the elastic axis and ranging from 1 kN/m to 10 kN/m in each direction (each symbol on the lines corresponds to a step of 1 kN/m)

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Fig. 17

Time series of blade tip torsion angle, uniform inflow, wind speed 11.4 m/s (rotational speed 12 rpm and pitch at 0 deg)

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Fig. 18

Time series of blade tip flapwise deflection, uniform inflow, wind speed of 11.4 m/s (rotational speed 12 rpm and pitch at 0 deg)

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Fig. 19

Centrifugal force effect on blade pitch due to large bending

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Fig. 20

Time series of blade root torsion moment, uniform inflow, wind speed of 11.4 m/s (rotational speed of 12 rpm and pitch at 0 deg)

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Fig. 21

Time series of blade root flapwise bending moment, uniform inflow, wind speed of 11.4 m/s (rotational speed of 12 rpm and pitch at 0 deg)

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Fig. 22

Time series of blade tip torsion angle, uniform inflow, wind speed of 18 m/s (rotational speed of 12.1 rpm and pitch at 15 deg)

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Fig. 23

Time series of blade tip flapwise deflection, uniform inflow, wind speed of 18 m/s (rotational speed of 12.1 rpm and pitch at 15 deg)

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Fig. 24

Time series of blade root torsion moment, uniform inflow, wind speed of 18 m/s (rotational speed of 12.1 rpm and pitch at 15 deg)

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Fig. 25

Time series of blade tip twist angle and root torsion moment, turbulent inflow, mean wind speed of 11.4 m/s, TI = 0.16

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Fig. 26

PSD of blade tip twist angle, turbulent inflow, mean wind speed of 11.4 m/s, TI = 0.16

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Fig. 27

PSD of blade root torsion moment, turbulent inflow, mean wind speed of 11.4 m/s, TI = 0.16

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Fig. 28

PSD of blade root edgewise bending, turbulent inflow, mean wind speed of 11.4 m/s, TI = 0.16

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