Increasing the operational capability of a horizontal axis wind turbine by its integration with a vertical axis wind turbine

Applied Energy 199 (2017) 479–494

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Increasing the operational capability of a horizontal axis wind turbine by its integration with a vertical axis wind turbine Bala Govind Department of Mechanical and Aerospace Engineering, North Carolina State University, Raleigh, North Carolina 27606, United States

h i g h l i g h t s

g r a p h i c a l a b s t r a c t


A new mechanism transfers torque

Sha Rotaon Reversal

CVT

from HAWT rotor to an integrated VAWT drive-train at high wind speeds.
Operational range is improved for prototype-scale 12 kW HAWT-10 kW VAWT combination.
A k-x (SST) turbulence model suggests safe rotor clearance for the integrated system.
Aerodynamic feasibility reveals effects of torque ripple on the combined power output.

a r t i c l e

i n f o

Article history: Received 31 January 2017 Received in revised form 8 April 2017 Accepted 26 April 2017

Keywords: Horizontal axis wind turbine Vertical axis wind turbine Hybrid wind turbine design Motion transfer Computational fluid dynamics Turbulence modelling

Primary Right Angle Transmission Shell-VAWT Moon Transfer VAWT Generator

a b s t r a c t A major difficulty encountered by a horizontal axis wind turbine is the limit of aerodynamic torque that it can withstand at high wind speeds. A novel strategy is proposed to improve the operational capability of a prototype scale system by increasing its rated wind speed for power generation. This is achieved by integrating its drivetrain with that of a vertical axis wind turbine supported on a common tower. Excess torque is transferred from the horizontal axis rotor to the vertical axis rotor’s drivetrain by coupling them using a continuously variable transmission. In this article, firstly, the concepts of motion transfer that facilitate this combined operation are discussed. A combination of a 12-kW horizontal axis rotor and a 10-kW vertical axis wind turbine is studied to estimate the increased benefit of increments in rated wind speed. Performance of this hybrid system is predicted at potential wind sites and is shown to exceed the standalone mechanical power output of both subsystems under different wind regimes. The critical criterion of the system’s aerodynamic feasibility is then investigated. Turbulence modelling is performed for a configuration which involves a combination of the NREL Phase VI rotor and a NACA 0021 profiled vertical axis H-rotor. A 3-D simulation, using a validated k-x (Shear Stress Transport) computational fluid dynamics model helps confirm the ability of both turbines to operate aerodynamically independent of each other. Further, by this methodology, a safe clearance between the two rotors is pre-determined. Analysis of turbulent flow scenarios reveals the characteristic effects of aerodynamic torque ripple experienced by the vertical axis wind turbine and its impact on combined power output. Parameters outlined in this article will be of assistance in the practical implementation of the integrated axes wind turbine. Ó 2017 Elsevier Ltd. All rights reserved.

E-mail addresses: [email protected], [email protected] http://dx.doi.org/10.1016/j.apenergy.2017.04.070 0306-2619/Ó 2017 Elsevier Ltd. All rights reserved.

HAWT Generator

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B. Govind / Applied Energy 199 (2017) 479–494

Nomenclature HAWT horizontal axis wind turbine VAWT vertical axis wind turbine CFD Computational Fluid Dynamics NREL National Renewable Energy Laboratory TSR Tip Speed Ratio CVT Continuously Variable Transmission RANS Reynolds Averaged Navier-Stokes (Equations) NACA National Advisory Committee for Aeronautics NASA National Aeronautics and Space Administration Rpm revolutions per minute Z 1 ,Z 2 ,Z 3 ,Z 4 tooth numbers of VAWT motion transfer gear-train r CVT ratio ucH cut in wind speed for the HAWT uHB rated wind speed for the HAWT rated wind speed for the VAWT uVB uFH furling wind speed for the HAWT uFV furling wind speed for the VAWT xH angular speed of HAWT rotor

1. Introduction Horizontal axis wind turbines (HAWTs) are the primary source of grid connected wind power [1]. A primary issue encountered is that at high wind speed nearing its wind rated speed, the HAWT’s operation is stall regulated or pitch controlled to limit power generation by reducing lift on its blades [2]. This limits the optimal torque exerted and causes the wastage of a portion of potential mechanical power at a site. Therefore, to better facilitate energy capture from the three-bladed turbine, drivetrain modifications and control schemes are being designed and improved to maximize the power drawn at varying wind speed. More recently, research is veering towards the incorporation of artificial neural networks [3,4] which form a framework of learning algorithms which help correlate on-site wind data with the power coefficient and pitch angles. In particular, the adaptive neuro-fuzzy inference system (ANFIS) [5–7] is gaining popularity. On a broader scale, it can also assist in maximizing profit from empirical wind profiles for wind farms [8–10]. Feathering and pitch actuation models are also being improved to optimize the angle of attack of the wind and the instantaneous Tip Speed Ratio (T.S.R) of the HAWT. These are supported by soft computing techniques for power coefficient optimization [11,12] and Maximum Power Point Tracking (MPPT) algorithms [13,14] which enhance the synchronous power output from the turbine’s grid-connected Permanent Magnet Synchronous Generator (PMSG) or Doubly-Fed induction Generator. Here, the use of the sliding mode control strategy [15–17] is a popular method which helps optimize the functioning of the turbine’s drivetrain elements. While these strategies may assist in compensating for the limitations in the existing capability of the HAWT, attention must pivot towards the fundamental design of the three-bladed system. To increase the mechanical power a HAWT can harness, changes in the design of the three-bladed rotor are being attempted. One imaginative method to draw more power is to access the stream of fast-flowing wind at higher altitudes by the crosswind motion of tethered wings [18,19]. Another bold approach entails radically increasing the swept area of the turbine. For example, a new design was tested by Vestas Wind Systems [20] wherein multiple 3-bladed rotors were supported at different eccentricities from a central tower to achieve a 900-kW output. While it is claimed that this reduces the Levelized Cost of Energy (LCOE) [21] and construction and transportation costs, the rotor

xHB xV saH scH saV scV R

q

k Cp Cq P  P1 A

v1

TKE

rated Angular speed of HAWT rotor angular speed of VAWT rotor aerodynamic torque on HAWT main shaft control torque exerted by HAWT generator aerodynamic torque on VAWT central shaft control torque exerted by VAWT generator radius of rotor density of air Tip Speed Ratio coefficient of power coefficient of torque gauge pressure projected area of rotor free stream wind speed Turbulent Kinetic Energy

overhang in such a configuration could be challenging. The key consideration for all such hybrid designs is that implementation at a megawatt-scale is governed by economy. This is compounded by the breakdown of critical components, thus increasing operational cost [22]. To improve the mechanical power output at the hub height of the HAWT, some improvements in the existing motion transfer mechanism within the nacelle have been undertaken. These include the use of direct-driven PMSGs which avoid the need of intermediate gearboxes. Also, to facilitate smooth gear ratio transition, the use of Continuously Variable Transmission [23–25] is being investigated. Contrary to a manual or a conventional automatic transmission, the operation of a CVT involves no torque interruption during change in angular speed. While CVTs have been investigated for active drivetrains in HAWTs to convert gusty wind power to stable alternating current for synchronous generation [26], practical designs which address optimum torque transfer and ability to raise rated wind speed haven’t been adequately implemented. With an apparently disparate purview, to harvest more energy near the ground surface, vertical axis wind turbines (VAWTs) too are being re-designed to up-scale their drivetrains. To save costs, intermediate gearboxes are avoided and instead, low solidity ratio models which facilitate low torque and high rpm central shaft rotation to optimize the use of alternators, are used [27]. Consequently, more effort has been directed to fabricate VAWTs of a 1–20 kW scale [28]. By arranging these turbines in arrays so they may mutually aid each other in rotation [29], it is claimed that they produce a higher aggregate of mechanical power. Unlike their horizontal axis counterparts, however, VAWTs cannot access the stream of fast flowing wind as they are located at a lower elevation, have a smaller swept area and suffer from turbulence at ground level. In an attempt to address these difficulties, the author proposes a new mechanism for a prototype scale wind turbine system. This incorporates the use of a shell-type VAWT installed on the same support tower as the HAWT. This constitutes a strategy to optimize the power output of the HAWT in a unique way. A novel motion transfer mechanism comes into operation when the wind speed at the hub height of the HAWT approaches or exceeds its original rated wind speed. The modified HAWT drivetrain is designed to transfer the excess torque exerted on the HAWT’s main shaft to the VAWT’s central shaft. This further utilizes a CVT for a purpose removed of its normal purpose of power transmission from rotor to

B. Govind / Applied Energy 199 (2017) 479–494

generator. As it is capable of smooth acceleration with no motion seizure, the CVT facilitates the seamless integration of the HAWT rotor drivetrain and the shell-VAWT drivetrain at high wind speeds. In this design, the excess mechanical power of the HAWT acts as a source and the VAWT generator acts as a sink. While two-ended torque transfer can be achieved until both turbines reach their rated wind speeds, a VAWT’s rotor may not normally facilitate the source of extra energy. The HAWT, therefore, possesses the higher angular momentum which must be limited. Since electric power is also independently generated by the VAWT at a lower height location on the same support tower, this arrangement provides a high-density power resource in wind farms which have space constraints. This progressive technology also aims to provide a failsafe option for operators to use a low-capacity aging HAWT at higher wind speeds without having to decrease aerodynamic torque by sub-optimally using feathering or pitch mechanisms. By reducing the loading on the HAWT’s original drivetrain, excess power is derived at the VAWT’s generator while also making optimum use of ‘real estate’ of the tower. This considerably increases the net swept area of the original turbine. The article is divided into two main sections which discuss the practical considerations of the proposed design’s implementation. The first section expounds on the design of drivetrain elements within the nacelle and H-rotor VAWT which benefit from torque transfer at hub-height wind speeds exceeding the standalone HAWT’s rated wind speed. It further offers the preferred mode of operation in different regimes. A combination of the NREL Phase VI rotor and a small manufacturer’s 10 kW VAWT and their relevant power characteristics is utilized to estimate of the mechanical power drawn for increments in new rated wind speeds. Power predictions at two potential wind sites in the state of North Carolina are also made. For instance, it is shown that during integrated operation, by preventing further reduction of power coefficient of a 12-kW HAWT beyond its original rated wind speed and using only a 1 m/s rise in rated wind speed, the system can yield an extra 2.9 kW at the VAWT’s generator. This is in addition to the 12 kW of independent HAWT power output and the 10–12 kW of independent VAWT output power. The second section critically examines the aerodynamic feasibility intended for the prototype scale hybrid system. It is important to minimize the possible mutual effects of turbulence caused by either turbine during combined operation. This is discussed using a 3-D time transient Computational Fluid Dynamics (CFD) approach to visualize the quasi steady-state performance of the integrated system. The hybrid design uses the NREL Phase VI S809 rotor and NACA 0021 profiled H-rotor geometry. The HAWT’s independent performance is first validated with NREL’s experimental results and the model is extended to study and confirm the aerodynamically independent operation of both turbine subsystems for cases of different angular speed and wind speed. The analysis further reveals the effects of the characteristic torque ripple experienced by a straight-bladed VAWT on the total mechanical power output. This will help design the H-rotor to operate in a pre-determined profitable range of angular speed. The stability of solutions of turbulence modelling also helps establish the suitability of the k  x (Shear Stress Transport) model [30,31] over the k  e (Realizable) model [32] for this investigation. An attempt to fabricate the design of the prototype scale integrated axes wind turbine will be undertaken shortly. Notably, the strategy of the proposed operational sequence of its drivetrain is generic. Only real-time data may facilitate an optimal learning framework for the control of the hybrid modes and the associated convertor circuity. This article will constitute an effort to explain

481

the concepts of motion transfer and more importantly, findings pertinent to aerodynamic feasibility of the hybrid system. The preferred method of actuation of drivetrain components is explained further in the author’s specification (US Pat. Appl. 2016/0201652 A1 [33]). 2. Primary drive-train and methodology to effect torque transfer from the horizontal-axis subsystem to the vertical-axis subsystem Under normal operating conditions, a conventional HAWT rotor facilitates a prime mover rotating at an angular speed conducive to independent synchronous electricity generation. When its main shaft’s angular speed dips below a reference rpm, auxiliary power is supplied to make the rotation synchronous. Alternatively, the pitch of its blades is varied to alter the Tip Speed Ratio (TSR) which normally lies in range of 5–6 for a 3- bladed turbine. In cases of excessive dynamic load, mechanical or electrical brakes may be applied to limit electricity generation until favourable conditions ensue. The energy supplied by the wind at every instant is proportional to swept area of the rotor and to the cube of the prevailing wind speed. Therefore, by limiting the force of the wind by suboptimally altering the Tip Speed Ratio, the full potential of the wind may, at higher wind speed, be unutilized due to limitations of generator loading. Instead, if there were a mechanism by which the excess mechanical power from this energy source at a higher altitude could be siphoned to a sink (or a lower energy source), then the HAWT rotor may continue operating at its rated angular speed in a free wind stream which exceeds its original rated wind speed. To obtain a profitable combined performance, the coupled system proposed must fulfil three criteria: (1) A mechanically robust motion transfer mechanism which ensures complementary rotary action between both subsystems. (2) Aerodynamic feasibility, i.e. the effect of the rotation of one turbine on the flow field should not adversely affect the other turbine. (3) An intelligent operation sequence/control algorithm which ensures a smooth transition in the hybrid power curve. The mechanism presented here serves to aid the original lowcapacity HAWT and accommodate the increment in extra torque at higher wind speeds. This is done by directing the extra aerodynamic torque exerted on the HAWT to the drivetrain of a VAWT mounted on the same support tower. This VAWT’s generator produces a control torque which may be varied electromechanically to track the optimal power point according to prevalent wind conditions. While this may avoid having to completely reinforce the blades and tower, local strength at certain divisions of the tower will need to be enhanced. From Fig. 1, it can be conceived that in the prototype-scale ‘Integrated Axes Wind Turbine’ system, a modified H-rotor VAWT is mounted on the support tower of a HAWT. The model for this prototype consists of two rotors whose design is contingent on a few trade-offs:
Drivetrain analysis [34] shows that the torque ripple effect common in VAWT designs due to upstream and downstream loading on VAWT blades is more pronounced for a 2-bladed design than a 3-bladed design. There would a time periodicity in the net torque acting on the VAWT shell and support tower if a 2-bladed design was used. This supports the selection of a giromill-type rotor for a simplified model to avoid complica-

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CVT

HAWT Generator

Primary Right Angle Transmission Shell-VAWT Moon Transfer VAWT Generator

Fig. 1. Primary drive train of the prototype-scale hybrid system consisting of horizontal axis wind turbine and vertical axis wind turbine sub-systems. Excess mechanical power is drawn at the VAWT generator by torque transfer.

tions arising from excessive superimposition of oscillations on the steady mean torque used to estimate the integrated axes turbine’s performance.
Exhaustive aerodynamic analysis [35,36], however, shows that VAWT designs with struts can introduce large amounts of drag and can significantly reduce its performance.
To alleviate ground turbulence effects the HAWT hub height may be increased but this necessitates more material for tower installation. To better visualize the sequence of motion transfer, the action of the new drivetrain is explained as a series of its components which now follows. 2.1. Rotary motion transfer within the vertical axis wind turbine and its active pitch mechanism The VAWT consists of a thin-walled shell which rotates about the central axis of the HAWT’s tower. Radial arms are attached by suitable means to the shell. This provision is made so that material is distributed at a greater radius from the axis and subsequently yields a greater moment of inertia. This generates greater torque and, therefore, more mechanical power. Animation sequence 1 (a) shows a variation in motion transfer arrangement from the tower-concentric rotating shell of the vertical axis turbine to the central shaft of the VAWT sub-system. As shown, the configuration consists an external ring bevel gear which transfers rotary motion to an internal central gear via bevel gear pinions. The pinions rotate faster that the main bevel gear. Noticeably, the apexes of all the bevel gears are coincident. It is possible to improve this arrangement. Animation sequence 1 (b) shows a face gear drive which is easier to assemble than the bevel gear drive. For instance, it is simple to equalize input and output velocities by selecting appropriate tooth numbers to fulfil the condition given by ZZ21  ZZ43 ¼ 1, where Z 1 and Z 2 are tooth numbers of gear and pinion for the outer drive and Z 4 and

Z 3 are tooth numbers of gear and pinion for the inner drive. However, installation of gear drives necessitates that the tower be made locally strong. If this modification is extended to a multi-megawatt utility, this may require a newer certification. Another implication of this alteration in a conventional VAWT’s design is the change in active pitch mechanism design. Its skeletal representation is shown in Animation sequence 2. Here, straight blades are mounted on the H-rotor shell by means of revolute joints at the ends of radial arms. In practice, the angular position of blades would depend on the elevation of a cylindrical slider whose position is in turn dependant on the action of a plurality of hydraulic cylinders which are housed outside the tower (refer specification [33]). This enhances serviceability and ease of replacement. 2.2. Continuously Variable Transmission for coupling the two turbine subsystems To resolve the issue of a smooth transition in rotary power transfer, an intermediate Continuously Variable Transmission unit is installed in the HAWT nacelle. Although numerous variants in CVT designs can be considered, the push-belt design is employed here for explanatory purposes. However, for a higher torque (100 kW–2 MW) wind turbine utility, the hydro-viscous type [23,24] is preferable. Fig. 2 shows an expanded schematic of this setup. In practice, compact actuation is possible. This configuration is governed by two sets of movable and fixed pulley sheaves and a segmented steel belt which transfers power between two parallel shafts akin a transaxle arrangement. The radii of the CVT’s sheaves depend on the varying speed ratio of the parallel shafts passing through them. The CVT’s transient ‘gear ratio’ is continuously varied by changing the width of the sheaves which result in seamless change in angular speed. As shown, the secondary main shaft’s excess torque acts as the drive. Unlike gear-trains, this ensures smooth coupling of the rotors without sudden changes in rpm of power transmitting elements of both supporting drivetrains. This further makes provision for a constant speed drive from the HAWT when the VAWT’s central shaft’s rpm is maintained constant after it reaches its rated value. This, however, requires that a highly durable steel belt of constant length be used to withstand the increased sheave pressure during compression. Fig. 2 also shows the mechanism of engagement with the CVT and is similar to the mechanism of a dog clutch. Hydraulic linear actuators engage toothed wheels to the slotted cavities in the sheaves of the CVT. Retractable telescopic shafts engage with the CVT via a spring mechanism. They are protected from wobbling effects by support fixtures. 2.3. Main Right-Angled Transmission within the HAWT nacelle The excess torque is transferred from the secondary main shaft to an eccentrically aligned horizontal shaft which is engaged with the Main Right-Angled Transmission. This transfer occurs only when the two parallel, horizontally aligned shafts are engaged with the CVT. There is an advantage to using two helically profiled bevel gears to transfer motion from the secondary main shaft the VAWT central shaft. They ensure that torque interplay between from the horizontal rotor to the central shaft occurs when the nacelle simultaneously yaws about the tower. This is shown in Animation Sequence 3. Here, the bevel gear which is integral to the nacelle rolls along the circumference of the other bevel gear which is integral to the central shaft. The latter’s axis of rotation corresponds to the axis of rotation of the nacelle and also to the longitudinal axis of the support tower of the HAWT. In the pre-

B. Govind / Applied Energy 199 (2017) 479–494

483

Fig. 2. Expanded schematic of mechanisms incorporated in (a) the HAWT’s nacelle, (b) the shell-VAWT’s rotary motion transfer and (c) simplified VAWT active pitch control.

ferred construction, it is desired that both bevel gears be equally sized to transmit uniform angular speed without speed reduction. 2.4. Shaft rotation reversal gearbox to ensure complementary rotary action One disadvantage of the push-belt CVT is that the belt is directional. It isn’t designed to transfer equal torque to the transaxle in the reverse direction. If the HAWT and VAWT rotors do not facilitate such complementary rotation by independent operation when engaged, there wouldn’t be profitable power transfer between the coupled drivetrains and as such, they would oppose each other. Therefore, for completeness in operating procedure, a gearbox containing a bevel gear clutch is installed prior to the engagement of the HAWT main shaft to the CVT in the nacelle. The action of this shaft rotation reversal mechanism in the drive train is shown in Animation sequence 4. As shown in Fig. 2, the bevel gear A is the driver. The secondary main shaft rotates in either clockwise or anticlockwise direction depending on which of the two gears B or C is engaged by the clutch. B and C rotate freely on the secondary main shaft. The clutch consists of a toothed collar and has a splined profile to engage with the output shaft. The gearbox further forms a secondary right angle transmission which transfers motion from the eccentrically aligned HAWT primary main shaft to the secondary main shaft. When actuated, this gear-

box ensures that the two parallel horizontal shafts engaged with the CVT complement the rotation of the VAWT central shaft. The integration of the aforementioned components embodies a new system that transmits optimal torque of horizontal and vertical axes rotors to their electricity generating components. Other support fixtures, fluid couplings and bearings [33] installed will mitigate problems caused by the fatigue loading and vibration in the tower. Arguably, to increase the rated windspeed of the HAWT rotor, a generator with a higher power rating could be installed, and thereby avoid the entire CVT/VAWT arrangement. However, it may only accommodate an increased aerodynamic torque which cannot far exceed the rated loading of the original drivetrain. The point of the design is to make the integrated components modular. While the idea of using a single generator in the nacelle to couple both turbines may be considered, it isn’t preferable for two reasons:
It is expensive to redesign or install a generator which was originally electrically rated to match the original HAWT drivetrain’s capabilities
Directly engaging an extended central shaft to the HAWT generator and without variable speed transmission may be mechanically and materially infeasible when scaled up. Further, during downtime or component service, operations of both turbines would be halted.

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3. Generic drivetrain coupling sequence and prediction of additional power derived The flow-of-control sequence leading to torque interplay between the horizontal and vertical axes rotors is shown in Fig. 3. The controller receives input data of wind speed at the HAWT hub height (uH ), wind speed at the VAWT mid-plane height (uV ), aerodynamic torque on the primary main shaft (saH ), aerodynamic torque on VAWT central shaft (saH ), angular speed of the secondary main shaft (xH ) and angular speed of the VAWT central shaft (xV ). The controller then monitors the gradient in angular speeds xH and xV . It may optimize independent synchronous generation from the HAWT according to the Maximum Power Point Tracking algorithm in the wind-speed range ucH 6 uH 6 uHB at the hub height of the HAWT, where ucH is the HAWT cut-in speed and uHB is the original rated wind-speed of the HAWT. It may further follow a similar tracking procedure or Hill Climb Search algorithm for the VAWT, i.e. in the range ucV 6 uV 6 uVB at the mid-plane height of the VAWT.

At every instant, the collective aerodynamic torque exerted on the HAWT rotor hub and blades is

saH ¼ 12 q  pR3  C q ðb; kÞ  uH ðtÞ2 .

Here, Tip Speed Ratio (T.S.R), k, equals

xH R uH

and the torque coeffi-

Simultaneously, a control torque, scH is exerted cient, C q , equals by the HAWT generator and is responsible for the horizontal axis rotor’s synchronous speed regulation. This control torque cannot exceed a maximum value scHB which may be pre-determined by the DFIG/PM generator characteristics. The most typical scenario of hybrid operation may arise as follows. If by site-specific design, the VAWT central shaft reaches its rated angular speed xVB before the HAWT reaches its rated rpm of xHB , xV would be maintained constant at xVB even before coupling the two drivetrains. Also, at the culmination of the mode of operation corresponding to ucH 6 uH 6 uHB , the HAWT’s angular speed would approach xHB and would need to be maintained constant by a combination of torque transfer and optimal pitch control. This implies that after coupling the two rotors, the controller would monitor the gradient of the rise in their rpm and by suitable actuation, vary the lift forces on the turbines’ blades by the pitch mechanism of either turbine to keep xH under Cp . k

Release HAWT and VAWT subsystems’ mechanical brakes

Check if starting torque(

) on HAWT is sufficient

Check if starting torque (

)on VAWT is sufficient

Yes

No

No

Yes

Vary HAWT blade pitch angle/ yaw angle/ supply auxiliary power

Monitor gradient of rise in secondary main shaft’s rpm by varying HAWT control torque

Vary VAWT blade pitch angle/ supply auxiliary power

Monitor gradient of rise in central shaft’s rpm by varying VAWT control torque

Check if secondary main shaft Yes

No

Check if rotational sense of both rotors would be complementary during engagement with CVT. Yes

No

Actuate Shaft Rotation Reversal Mechanism

Check if wind speed equals or exceeds

Yes No

Engage (or continue engagement of) HAWT secondary main shaft and VAWT central shaft by CVT

Vary VAWT generator’s to keep integrated HAWT’s under original rated

Continue HAWT’s independent operation

Disengage/ (continue disengagement of) HAWT main shaft and VAWT central shaft from CVT

Continue VAWT’s independent operation

Fig. 3. Mode of operation leading to the coupling of HAWT and VAWT drivetrains. It culminates torque transfer from the horizontal rotor to the vertical rotor’s drivetrain when uH exceeds uHB .

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xHB . Effectively, even though the wind-speed exceeds that corresponding to the HAWT’s rated angular speed, it continues to draw a larger magnitude of power due to the transfer of a fraction of torque which exceeds that of its independent rated operation. DP HAWT , the incremental power extracted, is a product of incremental torque DsH and the constant rated angular speed xHB . To implement this, the CVT must be brought into action to couple the HAWT’s main shaft and the central shaft of the VAWT generator at its desired constant speed of xVB . The additional power is optimized by varying VAWT control torque and keeping constant the angular velocity of both the HAWT main shaft and VAWT central shaft after their integration, while maintaining an optimal T.S. xV R. Ideally, CVT ratio r would be given by x . H The mode of hybrid operation is effective in the range uHB 6 uH 6 uFH , where uFH is the furling speed. Importantly, during this mode, effective torque on the HAWT during operation of the CVT equals saH  scHB . The excess torque can be expressed as

1 1 DsH ¼ q  pR3  C qH ðb; kÞ  uH ðtÞ2  q  pR3  C qHB ðb; kÞ  uHB ðtÞ2 2 2 ð1Þ C qH would be maintained constant at C qHB in this regime. If the transmission efficiency of the CVT and right angle transmission is say, gtr ; the equivalent torque transmitted to the VAWT generator is DscVH ¼



1 1 q  pR3  C qH ðb; kÞ  uH ðtÞ2  q  pR3  C qH ðb; kÞ  uHB ðtÞ2 2 2

 

gtr r ð2Þ

Consequently, the increment in mechanical power derived is equal to DscVH  xVBH . This is manifested as electromechanical conversion at the VAWT’s generator, subject to constraints of efficiency of transmission such as influence of slip and friction in the CVT variator. gtr should normally vary in the range of 70– 90%, depending on xH [37]. If uH subsides to a value less than uHB , the decoupling of the main HAWT shaft and the VAWT central shaft will follow. However, if uH exceeds the furling speed uFH of the integrated system, the HAWT’s further operation is stopped due to mechanical and electrical limitations. Feathering mechanisms and mechanical brakes would then be used for both rotors when uH P uHF and uV P uVF .

Table 1 Vital characteristics of SAWT Inc. PK-10 AB 10 kW (nominal rating) VAWT. Rated Power/Maximum Power Rated Wind Speed, uHB Cut-in Wind Speed, ucV /Cut-out Wind Speed, uFV H-rotor diameter Blade height Generator Type/Rated Voltage Speed Control

10 kW/12 kW 12 m/s 2 m/s/25 m/s 6m 6.2 m Permanent Magnet-3 Phase/AC 250 V Automatic Pitch Control and Hydraulic Brake

The two subsystems’ independent power curves are shown in Fig. 5. The independently operating HAWT’s uHB equals 12.5 m/s and the independently operating VAWT’s uVB equals 12 m/s. When uH exceeds uHB , torque transfer comes into effect. Fig. 5 further shows the consequent increase in mechanical power drawn from the integrated axes wind turbine. It far exceeds the combined individual capacity of both turbines. Predictions are also shown for increments of 1 m/s, 2 m/s and 3 m/s in uHB of the Phase VI rotor. The torque transfer to the VAWT generator in this range of uHB is immediately apparent. These predictions show that even if rotary motion is inefficiently transmitted by the CVT and the intermediate right angle transmission and HAWT’s rotation is maintained at a C qHB of only 0.524, or a C pH of 0.1, the integrated system would harness an additional 2.93 kW of mechanical power at the VAWT generator with an increase of only 1 m/s in uHB . This is in addition to the independent 10–12 kW mechanical power output of the VAWT. Further, this can be implemented without feathering or sub-optimally varying the HAWT’s blades’ pitch. This formulation will be used to draw parallels with the forthcoming aerodynamic feasibility discussion. To better estimate the performance of the integrated system at potential installation sites, wind profile data was collected from two onshore weather stations in the state of North Carolina. This data is archived by the State Climate Office of North Carolina at North Carolina State University. The data was collected for Hatteras, Dare County and Grandfather Mountain, Avery County and was sampled at an hourly rate for a 90-h cycle. By estimating the mechanical power available at the hub height of the HAWT

4. An illustrative power curve prediction As outlined in the preceding section, in its simplest mode of operation, the VAWT is designed to independently reach its rated mechanical power output before the HAWT reaches its independent rated mechanical power output. One preferred combination of a prototype setup entails integration of the two-bladed NREL Phase VI S809 HAWT rotor [38–40] and SAWT Inc.’s PK-10 AB VAWT [41,42]. The former has a blade length of 4.8 m and its rotor’s design will be discussed further in the forthcoming section of turbulence modelling of the hybrid turbine. The VAWT is a small manufacturer’s commercially available 10 kW H-rotor and is of comparable scale as the HAWT. Its details are listed in Table 1. Further, this VAWT’s capabilities are similar to those used in the forthcoming discussion on aerodynamic feasibility. Using Eq. (2) of the preceding formulation, the maximum excess torque exerted at the VAWT generator can be estimated. This varies with CVT ratio. Assuming ideal transmission, this increment in sH is shown in Fig. 4. Importantly, this predicts the degree of oversizing of the VAWT generator. This prediction is site-specific with a target of utilizing torque transfer at wind speeds exceeding the Phase VI rotor’s rated wind speed.

Additional Torque at VAWT Generator (in kN-m)

8 r = 0.5

7

6

r = 0.625

5

r = 0.75

4

r = 0.875

3

r = 1.0 r = 1.125 r = 1.25 r = 1.375 r = 1.5

2

1

0 14

14.5

15

15.5

16

Wind Speed u

16.5

17

17.5

18

(in m/s)

Fig. 4. Maximum additional design torque, assuming ideal transmission, exerted at the VAWT generator. This varies with CVT ratio r during the coupled operation of the NREL S809 rotor HAWT and SAWT Inc. PK-10 VAWT turbines and comes into effect when uH exceeds uHB .

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This far exceeds the Phase VI rotor’s performance even when the hybrid system is operated at a suboptimal C pH of 0.1. While this optimistic estimate assumes ideal transmission, future research into a robust control algorithm will better determine gtr derived from a friction model [43] of the CVT and intermediate rightangled transmission elements.

40

35

IAWT, u IAWT, u

30

IAWT, u

HB HB

HB

=17.3 m/s =16.3 m/s =15.3 m/s

Power (in kW)

25

5. A note on practical installation

20

15

NREL S809 HAWT SAWT Inc. PK-10 VAWT

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Fig. 5. Power curves showing mechanical power vs wind speed for (a) the independent operation of the NREL Phase VI rotor rotating at an enforced xH of 72 rpm and C pH 0.1 (b) independent operation of the SAWT Inc. PK-10 VAWT subsystem and (c) combined power output from the integrated system for modes of operation uH 6 uHB and uH P uHB , assuming ideal transmission for three plausible increments in uHB .

Here, an initial assessment of additional capital costs for components in the 20 kW HAWT – 10 kW VAWT integrated axis wind turbine prototype system is provided (Table 2), roughly in accordance with current market value. Realization of a small-scale system is made easier by increased expenditure on individual components to ensure a robust design. This may be further aided by aeroelastic analyses done by computational researchers for tower loading such as that for the NREL Phase VI rotor’s tower [44,45] and structural loading of a standard VAWT [46]. The tower design will be more challenging for scaled versions and models could be extended to predict dynamic response of tower modes for a multi-megawatt scale integrated axes wind turbine system. The increased breadth of the design relating to power electronics and spacing of cables and ladders inside a scaled tower lends to a future discussion. 6. Aerodynamic feasibility

(19 m) and the mid-blade height of the VAWT (5.2 m), the suitability of both sites can be gauged. The hourly wind profiles for the sites are shown in Figs. 6a and 7a. The mechanical power drawn by the VAWT subsystem approaches its independent rated performance at several instances during the cycle. The leading distinction is the predicted mechanical power output of the Phase VI rotor subsystem at the two locations. As shown in Fig. 6b, at Hatteras, Dare County, the original uHB is not encountered. The net output is, therefore, a simple sum of power of the two independently operating subsystems. However, the combined operation has decisive effect when uH exceeds uHB at the Grandfather Mountain site, Avery County where a wind speed uH up to 22.3 m/s is recorded. As shown in Fig. 7b, this furnishes an opportunity to design the integrated system for three increments in uHB . Assuming ideal transmission when the drivetrain is coupled, for a new uHB of 15.3 m/s, the system yields 28.8 kW while for a heightened uHB of 17.3, it may yield 33.2 kW.

It is critical to investigate the effect of turbulence caused by one turbine on the flow field of the other. Here, a computational fluid dynamics approach is used. The simulation justifies a minimum clearance between the rotors for pre-stall conditions. 6.1. Validation of the CFD model First, the accuracy of the commercial solver Fluent 16 is validated. The NREL Unsteady Aerodynamic Experiment (UAE) Phase VI provides a standard test case for 3D CFD HAWT rotor analysis. Experimental results are documented by the National Renewable Energy Laboratory. The test case was studied exhaustively by computational researchers [47–49] to validate their software and improve prediction of aerodynamic performance. The rotor geometry makes use of the NREL S809 airfoil, the geometry of which is modelled in GAMBIT. To implement a rotat-

25 u V at VAWT mid-plane height of 5.2 m

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Fig. 6b. Time-transient prediction of mechanical power from independent HAWT and VAWT subsystems and the integrated axes wind turbine (IAWT) for a1 m/s increment in uHB .

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45 IAWT uHB =15.3 m/s IAWT uHB =16.3 m/s IAWT uHB =17.3 m/s NREL Phase VI rotor SAWT Inc. PK-10 VAWT

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Fig. 7b. Time-transient prediction of mechanical power from independent HAWT and VAWT subsystems and the integrated axes wind turbine (IAWT) for three 1 m/s increments in uHB .

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Table 2 Estimated additional component costs for integrating a 10 kW VAWT on the common tower of a 20 kW HAWT (in k$). 10 kW VAWT Permanent Magnet generator and blade kit On-grid 3 phase inverter and control unit 19-meter Mono-pole kit (Support tower) (1700 kg) Push-belt Continuously Variable Transmission/Infinitely Variable Transmission HAWT right-angle transmission and shaft rotation reversal gearboxes VAWT shell motion transfer gear train Plurality of linear hydraulic actuators and housing unit for VAWT pitch mechanism Miscellaneous support members, bearings and cables

11.0 9.5 8.6 0.8 0.6 0.6 1.2 1.0

ing mesh methodology, sliding interfaces are defined between an inner rotating volume (containing the rotor) and a stationary outer volume representing the remaining fluid domain. The influence of the hub, nacelle and tower on the rotor aerodynamics is negligible, which is a fair approximation for an upwind turbine. The rotor radius is 5.029 m and the blades are assumed to be rigid. The theoretical definition of the S809 profile indicates that it has a very sharp trailing edge. Fig. 8a shows the development of the blade by the specified chord lengths, taper angles and twist angles. In this upwind configuration, the input wind speed is a constant 7 m/s with the nacelle oriented at a yaw angle of zero degrees. The rotor rotates at an angular speed of 72 rpm. Blade surfaces are defined as no-slip walls. Fig. 8b shows different boundaries and contrast in mesh resolution in the computational domain. The mesh is refined in the inner cylindrical region for better resolution near the rotor. Air density and viscosity are 1.225 kg/m3 and 1:78  105 kg/(m s) respectively. The mesh consists of 2,883,965 tetrahedral elements. The computations employ compressible Reynolds Averaged Navier-Stokes (RANS) equations, using the k  x (SST) model which is recommended for small-scale turbine systems [30,31]. The tip pitch angle is set to 3°. Unsteady inflow is neglected, considering that the experimental data set was derived from repeated measurements. A time step of 0.01 s is chosen for simulation. The parameter considered for validation is the pressure coefficient at varying span-lengths, given by

cp ¼

P  P1

1 2

q½v 21 þ ðxRÞ2 

ð3Þ

P  P1 is the gauge pressure of the point under consideration on the surface of the blade, v 1 is free stream velocity and R is the radial distance from the hub.

Fig. 8a. Development of NREL S809 airfoil by successively tapered and twisted profiles.

Fig. 8b. Computational domain and relative mesh resolution for validation of NREL S809 rotor’s aerodynamic performance.

Fig. 9 shows parity of CFD predictions with experimental results [47,50] after 2 revolutions. A maximum variance of 9% is observed. Therefore, this modelling is reliable and can be extended to reliably predict the aerodynamic performance of the integrated axes system. 6.2. CFD simulation of the integrated axes wind turbines’ operation 6.2.1. Modelling the hybrid system of the HAWT and VAWT The horizontal axis rotor chosen is the same 2-bladed NREL S809 rotor used in the solver’s validation. The HAWT’s nacelle is of 0.7 m  0.7 m in cross section and 1 m in length. The selection of a vertical axis rotor for an integrated system is specific design and depends on the predicted wind speed (and consequently, T.S.R). In this model, the rotor, concentric to the tower of the HAWT, has straight blades which are 4.8 m in height. Its blades have a uniform NACA 0021 profile with a chord length of 0.7 m. Its struts, six in number, are simplistically modelled and connect the air-foils to the rotating shell integral to the lower portion of the tower. Choosing an aspect ratio (diameter/height) of 2 yields a rotor radius of 4.8 m. The clearance lies between the tip of the lower blade of the S809 rotor at initial 6 o’ clock position, and the plane comprising the higher end surface of the NACA 0021 profiled blade of the vertical axis rotor. Again, the rotor zones are embedded in moving mesh regions and a stationary zone represents the remainder of the control volume. The computational domain is represented in Fig. 10a. A size function associated with the rotor’s solid surfaces effects a progressively increasing cell size. Fig. 10b shows the relative mesh resolution. A summary of boundary conditions may be found in Table 3. As a neutral atmosphere is assumed, the Boussinesq approximation is not employed to account for buoyancy effects caused by thermal variation. For all cases, the unsteady, time-transient simulation was performed with a time step of 0.01 s. A preliminary choice is made between two turbulence models. Case 1 (a) utilizes the k-x (Shear Stress Transport) equations while Case 1 (b) utilizes the k-e (Realizable) equations. Conditions entailing equal angular speeds for rotors of comparable moments of inertia and altitude-invariant wind profile of 7 m/s may be less realistic [51]. However, a higher angular speed of 72 rpm for both rotors and a lessened clearance of 4 m produce greater turbulence in the rotors’ clearance and this helps determine the degree of intermixing of streamlines of induced velocity in the near wake region and further interaction downstream. Upon establishing a convergent model, performance is analysed at two expected operating points. In effect, the same VAWT geometry is used but different angular speeds/power output are consid-

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Fig. 10a. Computational Domain for Cases 1 and 2.

Fig. 10b. Relative grid resolution in the computational domain.

ered. In Case 2 (a), the VAWT rotates at 45 rpm (CVT ratio r ¼ 0:625) and in Case 2 (b), a VAWT rotates at a lower speed of 36 rpm (CVT ratio r ¼ 0:5). The HAWT has an increase of 2.4 m in hub height, while its speed is a steady 72 rpm. The free stream wind speed is increased to 9 m/s. As the choice of a vertical axis rotor is dependent on site location and requirement to integrate with a micro-scale grid, an independent power rating range of 5– 10 kW for the vertical rotor is targeted.

6.3. Results and discussion of turbulence modelling of the integrated axes wind turbine A neat convergence at each time-step is achieved after approximately 25 iterations. A qualitative analysis of the simulation of the rotors’ combined operation in the flow field now follows. The parameter used for visualization of effects of vorticity is Turbulence kinetic energy (TKE). TKE is the mean kinetic energy

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Table 3 Summary of Cases 1 and 2. Case 1 (a) Turbulent Model used Boundary Conditions

Domain Size (x, y and z dimension) Mesh strength

Case 1 (b)

k  x (Shear Stress Transport) k  e (Realizable)
v 1 = 7 m/s
Rotor clearance of 4 m
xH ; xV = 72 rpm 117 m  49 m  80 m 2,583,400 polyhedral cells (varying resolution)

per unit mass associated with eddies in turbulent flow and is characterised by measured root-mean-square (RMS) velocity fluctuations. TKE can be produced by fluid shear, friction or buoyancy, or through external forcing at low-frequency Eddie scales [51]. This energy is transferred down the turbulence energy cascade, and is dissipated by viscous forces at the Kolmogorov scale. In the integrated axes wind turbine model, TKE is measured in relation to the centre-lines of three subsystems extending through the length (x-positions) of the computational domain and which lie in the x-y plane: (1) The HAWT subsystem whose equatorial centre-line passes through the origin of the domain and coincidentally, the centre of the nacelle. (2) The rotor clearance, the equatorial line of which passes though the tower at a point 7 m below the nacelle for Cases 1 (a) and (b) and 8.2 m below the nacelle for Cases 2 (a) and (b). (3) The VAWT subsystem whose equatorial centre-line passes through a point 11.4 m below the nacelle, for Cases 1 (a) and (b) and 13.8 m below the nacelle for Cases 2 (a) and (b). Fig. 11 shows the variation of Turbulent Kinetic Energy along these centre-lines for two instances of flow time for Cases 1 (a) and (b). The simulation also helps visualize if the streamlines indicate a mutual interaction in the near and far wake regions of the hybrid system. The wake development is shown by means of velocity contours plotted on iso-surfaces spaced at consecutive distances of 1 m in the computational domain. An initial interaction of the two near wakes is manifested as an intermixing and circulation of velocity vectors. This circulation is a temporary phenomenon which is transported downstream and diminishes with progressing time. The near wake behaves akin a body of revolution reducing the effect of the free stream which diverges around it and converges again. The tower shadow effect does not bear significant aerodynamic influence. As Fig. 11 shows, both k  x (SST) and k  e (Realizable) models predict similar trends in far wake regions. The k  e (Realizable) model, however, has stability issues due to numerical stiffness when flow is separated. The k  e model performs poorly for the complex flow involving steep pressure gradient and change in streamline curvature near the tower. On the other hand, the k  x equation (SST) model more accurately predicts vorticity at small magnitudes and time scales and proves convenient near blade surfaces. It is, therefore, selected for pre-stall conditions in Cases 2 (a) and (b). The development of downstream-wake region shows that as time progresses, there is lessened effect by either turbine on the mean wind speed in the intermediate clearance (IAWT centreline). The only major variance in vorticity and TKE is only encountered at the position of the tower which presents an obstacle (xposition equals zero). The TKE is found to have significantly reduced after 10 revolutions of both turbines. Visibly, after a quasi-steady flow regime is reached, the circulation caused by

Case 2 (a)

Case 2 (b)

k  x (Shear Stress Transport)
v 1 = 9 m/s
Rotor clearance of 6.4 m
xH = 72 rpm
xH = 72 rpm
xV = 45 rpm
xV = 36 rpm 2,640,669 polyhedral cells (varying resolution)

either rotor does not appreciably affect the other’s operation for a free stream speed of 7 m/s. In Cases 2 (a) and (b) which use the preferred k  x (SST) model, an expanded wake is formed behind both rotors in their region of induced velocity. Again, the initial low-pressure circulation is not sustained and is carried downstream. Fig. 12 indicates that after 15 s of flow time, a quasi-steady flow regime is established, albeit weak intermixing in the far wake region. With an increased wind speed of 9 m/s, an increase in 2.4 m in clearance permits the HAWT rotor to operate under little influence of the VAWT rotor and this may be considered a safe configuration. 6.4. Mechanical power estimation and effects of VAWT torque ripple on the combined operation Having established the hybrid system’s aerodynamic feasibility, a straightforward prediction of instantaneous mechanical power from each rotor can now be made. A product of the aerodynamic torque on each turbine about their relevant axes, as reported by the validated solver, and their respective angular speeds, i.e. 7.53 rad/s for the S809 rotor and 4.71 rad/s (Case 2 (a)) and 3.76 rad/s (Case 2 (b)) for the NACA 0021 rotor, is taken. This output, relative to flow time is shown in Fig. 13(a) and (b). Interestingly, when a quasi-steady flow regime is achieved, the time-response of mechanical power cases reveal that HAWT’s performance is relatively unaffected by the VAWT’s rotation when given a rotor clearance of 6.4 m. This is evident from Fig. 13 (a) and (b). Further, its independent output in the neighbourhood of 11 kW correlates reasonably with experimental and predicted power curves [40] of the Unsteady Aerodynamics Experiment Phase VI. The violent variation in axial moment exerted on the VAWT typical of this H-rotor type design will also cause its characteristic torque ripple effect [52] on its generator during its practical realization. This is the natural outcome of integrating torque produced by aerodynamic loads due to its rotating blades’ interaction with steady wind. Case 2 (b) is instructive as it shows that enforcing xV at 36 rpm is hardly profitable (see Fig. 13(b)). It effects an averaged drop of 8 kW (40 percent drop in efficiency) in net output of the integrated system’s output from that obtained at xV of 45 rpm in Case 2 (a) (see Fig. 13(a)). The compensated torque per cycle shows that energy is not harvested for a portion of its rotation and for a small period, the VAWT behaves akin a propeller that adds energy to the free stream. Also, after12 s of flow time, net saV drops sharply to produce a near –zero (a mean of 0.13 kN m) positive moment. 7. Conclusions and scope of future work The design of an integrated axes wind turbine has been proposed and discussed. Its novel drivetrain involves the coupling of a horizontal axis wind turbine and a vertical axis wind turbine at

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0.6 0.6 HAWT center -line IAWT center-line VAWT center-line

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Fig. 11. Time-transient wake development for a rotor clearance of 4 m (v 1 = 7 m/s) at instances of 6 s and 10 s of flow time. Above: Velocity contours plotted on iso-surfaces separated by 1 m in the computational domain. Below: Turbulent Kinetic Energy comparison along centre-lines through the length of the control volume (x-position) for k  x (SST) model (Case 1 (a)) and k  e (Realizable) model (Case 1 (b)). The latter has stability issues.

high wind speeds. This facility enhances the existing capability of a prototype scale HAWT while simultaneously producing power from a tower-mounted VAWT. Further, it provides a source of high-density wind energy conversion. A generic mode of operation to facilitate the torque transfer from the horizontal-axis rotor to the drivetrain of the vertical–axis rotor has also been presented. To investigate its mechanical power capabilities, two potential wind sites were selected and the output for plausible increments

in rated wind speed was estimated. The change in the hybrid power curve reveals that the coupled system would yield power far exceeding the individual capacity of either turbine subsystem. Predictions for the small-scale system show, for example, a potential nominal power of up to 33.5 kW for a 3 m/s increase in rated wind speed when the S809 rotor HAWT is sustained at even a low power coefficient of 0.1 after it reaches independent rated power output.

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Fig. 12. Time-transient wake development for a rotor clearance of 6.4 m and v 1 = 9 m/s, using a k  x (SST) model. Above: Turbulent Kinetic Energy comparison along centre-lines through the length of the control volume (x-position) for Case 2 (a). Below: Turbulent Kinetic Energy comparison along centre-lines through the length of the control volume (x-position) for Case 2 (b) for two instances (7.5 s and 15 s) of flow time.

Turbulence modelling of the combined operation was used to confirm the aerodynamically independent operation of the NREL Phase VI rotor and a small manufacturer’s 10 kW VAWT used in combination. Permutations of wind speed conditions and angular speed of the two rotors assist in establishing safe configurations with varying hub height by using Turbulent Kinetic Energy as a parameter to find a safe rotor clearance. The model also helps pre-determine the aerodynamic torque on both rotors in a quasisteady regime. The k  x (SST) model proved more stable for pre-stall conditions. It allays concerns of possible mutual influence of either rotors’ rotation on the other’s torque production. In practice, it would behove the designer to consider an accurate windspeed profile from the ground up while implementing the recommended system. The effect of torque ripple is predominant. As suspected, not all angular speed ratios are profitable. In one case, a decrease of only 12 percent in the H-rotor VAWT’s angular speed causes an overall decrease of 8 kW of the integrated system during the two rotors’ normal, aerodynamically independent operation. With the current design, it is indicated that the independent operation of both rotors is relatively unaffected by their superimposition on the same tower. This provides a guideline for designing the rotating shell’s pitch mechanism to maintain its angular speed above predetermined thresholds.

Future work would include examining the cyclical variation in vertical axis rotor’s speed to determine the sensitivity of a suitable controller. If used in a scaled version, strength of the mechanical linkage to the generator depends on such harmonic aerodynamic torque fluctuations. Here, the smoothening of transient response caused by torque ripple by the VAWT is necessary for synchronous power generation from the rotors and their ongrid integration. Further, a complete friction model will be formulated to accurately determine the losses in transmission in the coupled drivetrain and CVT. A practical fabrication of the prototype scale integrated system will be undertaken shortly. This will determine, by extensive testing, the suitable sliding-mode control logic for optimal power point tracking of hybrid modes for various Tip Speed Ratios. While simulation implies that small-scale, standard rotors and blade profiles for a prototype system are profitable, the concept should bear fruition in a multi-MW setting and could re-power aging HAWTs using scaled Darrieus VAWT designs. Objectively, this hybrid design would greatly increase net swept area of a conventional wind turbine and is assisted by the two-ended torque transfer that helps increase rated wind speed.

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Appendix B. Supplementary material

30 VAWT (NACA 0021) subsystem-45 rpm HAWT (NREL S809) subsystem-72 rpm Integrated axes wind turbine

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Supplementary data associated with this article can be found, in the online version, at http://dx.doi.org/10.1016/j.apenergy.2017. 04.070.

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(b) Fig. 13. Transient predicted mechanical power for HAWT, VAWT subsystems and integrated axes system when xH = 72 rpm. (a) Profitable combined output when xV = 45 rpm. (b) Near zero independent VAWT contribution when xV = 36 rpm.

Acknowledgements The author would like to acknowledge the discussions and provision of computational resources at North Carolina State University, Raleigh and help with explanatory sequences by Er. Thắng Nguyễn, Hanoi. Appendix A. Explanatory Animation Sequences Sequences 1 (a) and (b) Sequence 2 Sequence 3 Sequence 4

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Shell-VAWT’s motion transfer mechanism Shell-VAWT’s pitch mechanism Simultaneous yawing and motion transfer Simplified shaft rotation reversal mechanism

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