Power Collection from Multiple HydrokineticGenerators Utilizing Advanced Control

Robert J. Cavagnaro∗, Benjamin Strom†, Brian Polagye† and Andy Stewart∗∗Applied Physics Laboratory

University of Washington, Seattle, WA 98105E-mail: rcav@uw.edu, andy@apl.washington.edu

†Mechanical EngineeringUniversity of Washington Seattle, WA 98105E-mail: strombw@uw.edu, bpolagye@uw.edu

Abstract—A common DC bus power collection scheme is con-sidered for a two-unit array of cross-flow hydrokinetic turbinesutilizing Advanced control schemes. These controllers increasemechanical rotor efficiency while introducing large, transientspikes of bidirectional current. A numerical simulation in theSimscape Power Systems environment driven by results fromlaboratory-scale turbine experiments is constructed to evaluatesystem efficiency from mechanical power input to active powerdelivered to the grid. The combined effects of advanced controland power collection result in a 13% system efficiency increaseover a baseline case, while the collection scheme alone resultsin a 5% increase. Generator losses are determined to mostsignificantly impact system performance. Results of this studyare intended to inform the completion of an electromechanicalemulation machine to further evaluate generators and powerelectronics specified for field deployment. Allowances for heatdissipation and the interconnection of units are anticipated to bemajor engineering challenges that must be overcome for reliable,cost-effective implementation once deployed at sea.

Index Terms—Power collection, advanced control, turbines,simulation, grid integration

I. INTRODUCTION

Colocation of hydrokinetic generators in arrays necessitatesthe planning and design of subsea power collection, condition-ing, and transmission systems. Similar to considerations forthe development of offshore wind arrays, efficiency, reliability,and capital cost are to be weighed in the selection of anoptimal electrical interconnection system [1]. Additionally,highly variable power output due to spatial and temporal fluc-tuations in currents and waves over relatively small distancesand time scales necessitates a high degree of control andcoordination of individual generators. Finally, control schemesare being developed requiring intermittent motoring (activepower consumption) or reactive power consumption to achievehigher mechanical conversion efficiency for individual devices[2], [3]. However, the electrical efficiency of such controlstrategies is unclear.

Cross-flow turbines, for which the axis of rotation is per-pendicular to the direction of flow, are receiving renewedinterest due to several advantages when applied to energy cap-ture from flowing water, despite the complex hydrodynamicsinvolved in their operation. These include lower maximumblade speed, omni-directional energy harvesting while in a

vertical orientation, and suitability for tight array packing. Asmost designs do not allow for control through varying thepitch angle of individual blades, as is common in axial-flowwind and hydrokinetic designs, rotor speed control by varyinggenerator torque is implemented. Coupled with a synchronousor permanent magnet generator, electrical power output fromthese devices will be variable in both voltage and frequencyif the system is expected to maintain peak efficiency throughfluctuations (both short time-scale, such as turbulence and longtime-scale, such as progression of a tidal cycle) of inflowconditions. Therefore conditioning for grid integration willmost likely be required.

Many studies have been conducted analyzing various pro-posed topologies and technologies for distributed generationand microgrids. Specifically on the topic of power collectionfrom distributed sources, options for coupling broadly includeinterconnection on a DC bus, conversion to and collection asAC power (i.e., grid) frequency, at higher AC frequencies (typ-ically for cases where high-frequency AC loads are present),and hybrid AC/DC layouts with separate buses [4]. Losses inthese systems accumulate with additional layers of conversion;choosing a topology best suited to the nature of both generatorsand loads is critical for high efficiency [5].

A method of power collection with all collocated devicesconnected to individual bidirectional AC/DC converters ona low-voltage DC bus as depicted in Fig. 1, is proposedas a solution to efficiently and cost-effectively accommodateadvanced variable-speed and coordinated control schemes forcross-flow turbines. An initial evaluation of the electricalefficiency of turbine generators under these controllers andpower collection design is conducted in simulation utilizingthe Simulink Simscape Power Systems block set and sim-ulation environment [6]. The hypothesis that overall systemefficiency can be improved with decreased power spikes bycoordinating the power consumption phase of one turbinewith the power production phase of a co-located second whileboth are coupled to the same DC bus is tested. This studywill be used to inform development of and form a basisof comparison for a similar study on a hardware-in-the-loopelectromechanical emulation machine (EEM), followed byprototype testing at demonstration scale. The EEM will consist

Fig. 1. Power collector schematic for pair of turbines on a common DC bus

of the actual generators and power electronics to be usedin the demonstration deployment. Additionally, considerationsfor the marinisation and deployment of the demonstrationof the proposed power collector and associated transmissioncomponents are discussed.

II. METHODS

The presented simulation of hydrokinetic generation en-compasses conversion of input mechanical power to outputelectrical power. Components of the electrical system andsimulation parameters are detailed following descriptions ofthe turbines and control schemes under test and how inputdriving the simulation is derived from experimental results andappropriately scaled.

A. Cross-Flow Turbines

Significant challenges exist in the creation of accurateand robust numerical methodologies for studying cross-flowturbine performance due to the complex hydrodynamics of therotors. Small, laboratory-scale models are often constructedand tested to gain insight into designs otherwise unachievablewithout building a full-scale prototype. In addition to theadvantage of lower cost compared to a full-scale model, theseoffer the benefits of testing in a laboratory environment whereinstrumentation and monitoring are more easily obtained andallowed. Two identical rotors of 17.2 cm diameter (D) and24.3 cm height (H) are constructed from two symmetric,straight NACA0018 blades for the purpose of evaluatingperformance and testing advanced control strategies.

The rotors, shown in Fig. 2, are designated ‘upstream’ and‘downstream’ when used in combination as members of anarray based on their relative streamwise positions in the flow.Both rotors are mounted on central shafts while blades areterminated at circular endplates. The upstream turbine’s shaftis mounted directly to the face of a servomotor, while thedownstream turbine’s shaft is supported by a bearing on afixed plate and coupled to its servomotor at its upper end. Six-axis load cells (upper and lower for downstream, only upperfor upstream) measure reaction forces, notably torque (τ )imparted to and from the rotor. The servomotors and associatedmotor drives can run in both directions using speed or torquecommands and as both motors and generators (4-quadrant

Servo motorLoad cell

Load cell

TurbineRotor

Bearing

U∞

Fig. 2. Laboratory-scale cross-flow turbine experimental setup

operation). Position and angular velocity are measured usingencoders internal to the servomotors with a resolution of10E+6 counts per revolution. Tests are run in a recirculatingflume, providing inflow at 0.7 m/s, measured with an acousticDoppler velocimeter at a distance of 5D upstream of theupstream rotor. Additional detail on the turbine and testapparatus can be found in Strom et al. [7].

A widely-used performance metric for turbine mechanicalefficiency is the ratio of mechanical power to kinetic powerof the inflow, the coefficient of performance,

CP =τω

12ρAU

3∞

(1)

where ω is rotation rate in rad/s, ρ is the fluid density, A is theprojected frontal area (DH), and U∞ is the reference inflowvelocity. Performance is a function of the turbine’s tip-speedratio,

λ =ωR

U∞(2)

in which R is the radius of the rotor. Evaluated in isolation,maximum CP under constant speed control for both turbineswas 0.26 at λ of 1.2 with constant 0.7 m/s inflow.

B. Evaluated Advanced Control Scheme

An advanced control strategy is currently being developedfor cross-flow hydrokinetic turbines to emulate aspects ofactive pitch control while simultaneously taking advantage oftheir ability to be tightly packed in groups. The method is ahybrid of two concepts: intracycle angular velocity control andcoordinated control for arrays. The former involves varying therotation rate of a rotor as a function of blade position in orderto favorably change the blades’ nominal angle of attack withinthe period of a single revolution. Sinusoidal and semi-arbitraryangular velocity profiles are optimized for maximum CP

(calculated over complete revolutions) by varying amplitudeand phase parameters. Depending on the phase of rotation,

TABLE ILABORATORY AND FIELD-SCALE TURBINE PARAMETERS

Lab-Scale Field-ScaleD (m) 0.172 0.857H (m) 0.234 1.168A (m2) 0.040 1.001U∞ (m/s) 0.7 2.0

the turbine’s servo acts as both motor and generator twice(once each per blade) per revolution. Therefore, under theassumption of stable inflow conditions, the electrical poweroutput of the system is expected to have transient peaks ofboth positive (to the grid) and negative (from the grid) powerat a frequency of twice the angular rotation rate. Coordinatedcontrol, developed for arrays of closely-spaced cross-flowturbines, enables predictable and beneficial interaction of adownstream turbine in another turbine’s wake. Optimization isperformed for both relative spacing of a pair of turbines andcoordination of their phases in time, such that CP for the arrayis maximized. The studied hybrid method utilizes an optimalcross-stream separation of 0.8D at a fixed streamwise sepa-ration of 1.5D between upstream and downstream rotors andenforces phase coordination while optimizing both turbines’angular velocity as per sinusoidal intracycle control. Furtherdetail and discussion of these controllers and the experimentsproducing the results referred to herein, including design andoptimization of the two-turbine array are presented in Stromet al. in the same proceedings [7].

C. Scaling to Target System

Laboratory-scale turbines for which these controllers areoptimized for are roughly 1

25 of the projected frontal area and15 of the diameter of similar field-scale turbines in develop-ment. Relevant turbine parameters for both scales are given inTable I. As the power collection hardware being consideredis intended for the field-scale versions, input to the electricalsystem simulation representing the turbines under advancedcontrol must be appropriately scaled. Simulation inputs aretime series of ω and τ from experimental conditions yieldingthe highest array CP while performing coordinated intracycleangular velocity control. A trace of this input for one fullrotation of both upstream and downstream turbines is shownin Fig. 3.

Low rotation rate and high torque, inherent to the cross-flowturbines studied, is suboptimal for standard commercial off-the-shelf electrical machines. Therefore, field deployments andsimulation include a 16:1 gearbox between the main turbineshaft and generator. Due to dependence of turbine performanceon λ, angular velocity command is scaled such that theinstantaneous magnitude of λ remains constant between scales,

ωf =2NλlU∞,f

Df(3)

with subscripts f and l denoting field and lab quantities,respectively, and N indicating the gearbox ratio. As ωf

becomes lower in amplitude, it must also be stretched in time

0 0.1 0.2 0.3 0.4 0.5 0.60

100

200

300

Sp

eed

(rp

m)

US TurbineDS Turbine

0 0.1 0.2 0.3 0.4 0.5 0.6Time (s)

-2

0

2

4

To

rqu

e (N

m)

US TurbineDS Turbine

Fig. 3. Laboratory-scale angular velocity and torque for upstream/downstreamturbines

to maintain similarity of phase between scales such that, forthe same number of complete turbine rotations,

1

N

∫ωfdt =

∫ωldt. (4)

Practically, this is accomplished by resampling the ωf timeseries to be proportionally longer due to slower rotation by afactor of ωf

ωl. Time series of instantaneous τl are scaled such

that instantaneous CP is maintained between scales, yielding,

τf =τlωlAfU

3∞,f

NωfAlU3∞,l

(5)

and is similarly stretched in time. A moving average smooth-ing filter with a window size of 20 points (data collected at 1kHz) is applied to the torque command to improve simulationstability. In simulation, a ‘full’ rotor refers to a turbine withlisted field dimensions, while a ‘half’ rotor refers to a turbinehalf the height, but same diameter as the field dimensions. Thisis intended to halve the magnitude of input torque withoutaffecting speed commands, effectively halving input power.

D. Power Collector Simulation

Power collection on a common DC bus was selected toaccommodate the needs of coordinated intracycle angular ve-locity control for cross-flow turbines. This layout was selectedfor analysis, as it minimizes the number of power conversionstages while accommodating instantaneous power productionand draw; energy exchange occurs on the DC bus wheremultiple linked generator drives are connected. This removesindividual inversion and synchronization stages, replacingthem with a single higher capacity inverter maintaining gridvoltage and frequency.

Previous studies have utilized Simscape Power Systemsto model the interactions of multiple hydrokinetic generators

TABLE IISIMULATION PARAMETERS

SimulationDesignation # Turbines Turbine

Size Control Type

1 2 HalfCoordinated intracycleangular velocity,upstream/downstream

2 2 Half Intracycle angular velocity,upstream/upstream

3 1 Full Intracycle angular velocity,upstream

4 1 Full Baseline constant speed

with a utility grid, augmenting the body of work on similarsystems for wind energy conversion, including hybrid wind-solar arrays [8]–[10]. However, a study evaluating multiplediscrete generators utilizing advanced control and with hightemporal resolution appears absent in literature. Additionally,the simulation environment allows turbine parameters andcontrollers initially developed at small-scale to be evaluatedand iterated at high speed and with relative ease compared tophysical experimentation.

Simulations are conducted for four scenarios in order todifferentiate the effects of control and power collection on sys-tem performance. First, two half-rotor, field-scale turbines on acommon DC bus are simulated with scaled input for upstreamand downstream rotors running coordinated intracycle angularvelocity control. Second, the same simulation is conducted, butboth rotors are commanded to enact the upstream turbine’sspeed and torque commands. This is implemented to deter-mine the contribution of coordination between and mechanicaloptimization of upstream and downstream rotors to systemefficiency. Third, a single, full-rotor field-scale turbine is sim-ulated with inputs for the upstream rotor. By comparing resultsto those of second test, the contribution to system efficiency ofsplit generation and subsequent collection can be determined,as both runs employ equal input mechanical power. Lastly,one full-rotor field-scale turbine is simulated with a simpleconstant speed control, set for optimal performance at nominalfield-scale operating inflow velocity. This run is intended forcomparison of the advanced controllers with baseline turbine,generator, and system efficiencies. Efficiency calculations forrotors use a constant, nominal inflow velocity over completeturbine rotations. Specifications for the four simulation topolo-gies are summarized in Table II.

The Simscape Power Systems model of the common DC buspower collector consists of four major submodels: permanentmagnet synchronous machines (PMSM), AC vector controllerswith speed regulators and inverters, a bidirectional voltagesourced converter (VSC) maintaining a stable DC link, anda utility grid model to source, sink, and transmit power. Eachof these subsystems is adapted from the Simscape PowerSystems Specialized Technology library and example code [6].An image of the top-level simulation topology is shown in Fig.4.

Generators to be used as the power take-off (PTO) of thefull field-scale turbines under advanced control are Siemens

TABLE IIIGENERATOR SPECIFICATIONS FOR SIMULATION

SimulationDesignations 3 1, 2 4

# Phases 3 3 3# Pole pairs 4 4 4

Torque constant(Nm/A) 2.37 1.152 1.013

Voltage constant(VLL/krpm) 286.6 139.3 122.5

Flux linkage(V.s) 0.395 0.192 0.169

Stator phase resist.(Ω) 0.175 0.085 0.092

Armature inductance(H) 4.5E-3 9.5E-4 9.8E-4

1FK7105-2AC7 servomotors, modeled in simulation by spec-ifying nameplate data in the PMSM block. As torque com-mands for half-rotor simulations are half of full-scale, asimilar PMSM is specified for advanced controllers, but ratedfor roughly half power output. Additionally, an alternative,lower power rating PTO is specified for simulation runs ofthe full rotor under constant speed control. Torque limitsare assigned as nameplate values where available and arerespected throughout simulation. Generator specifications aresummarized in Table III. All generators are assigned nom-inally low rotational moments of inertia of 0.001 kg-m2.This simplification allows the arbitrary assignment of bothspeed and torque to the generators, allowing optimized inputsexperimentally generated at laboratory-scale to be applied atfield-scale power levels.

Closed-loop speed control is utilized in all simulations andachieved with the combination of a proportional plus integral(PI) and vector controller. The PI loop compares generatorspeed to input setpoint, setting appropriate ramp rates andlimits, and results in an electrical torque command. The vectorcontroller uses the specified PMSM parameters and electricalphase angle to convert electrical torque to d/q-axis current,compares this to measured current, and generates inverterpulses at a maximum frequency of 50 kHz to switch IGBTs onwhen current error is outside the bounds of a hysteresis bandof +/-0.05 A [6]. Inverters are set with snubber resistance andcapacitance of 5 kΩ and 1E-6 F respectively, with 0.2E-3 Ωon-state resistance. PI gains are tuned for tight tracking of thespeed input, such that for the simulation results presented,maximum absolute speed error does not exceed +/-3% ofcommanded value. Torque limits are set to the maximumallowable specified by the generator manufacturer of +/-150Nm for the full-scale generator and +/-75 Nm for the half-scalegenerator.

Both the DC link and AC feed are regulated with a VSC.The VSC allows bidirectional power transfer by controllingthe magnitude and sign of DC link current. Additionally, itmaintains unity power factor, frequency at 60 Hz, and levelof 240 V on the AC side of an IGBT bridge converter usingpulse-width modulation at a carrier frequency of 1.98 kHz [6].DC voltage is controlled with a PI loop to enact a setpoint of

Fig. 4. Power collector simulation overview

600 V. Gains are tuned to maintain constant voltage to within+/-1% of the nominal value. The VSC inverter is modeled withsnubber resistance of 1E+6 Ω and 0.2E-3 Ω on-state resistance.

A 240 V / 25 kV transformer couples the generation systemto the transmission and grid model. It has magnetizationresistance and inductance of 500 per-unit (pu) referred to itsnominal power rating. This rating is adjusted depending on thepeak power observed in initial simulations and ranges from 5kVA to 30 kVA. The grid model consists of 5 km and 14km 25 kV feeders, 2 MW and 30 MW active power loads,and a 2 Mvar inductive load, fed from a 120 kV three phasesource with an internal impedance ratio (X/R) of 7 [6]. Three-phase, positive sequence active power to and from the grid ismeasured at the coupling transformer.

Simulation starts with an initialization phase, in which theDC link is charged to its nominal value. Constant speed andtorque inputs are applied to bring the electrical machines upto their initial speeds. Variable velocity and torque commandsare then applied for 2.5 seconds, allowing for over 2 completeturbine rotations. The simulation is run with fixed, discretetime steps of 1E-6 s for accuracy and high resolution results.

System performance is quantified at several steps frommechanical input to electrical output: mechanical conversionefficiency (CP ), electrical efficiency evaluated at the DC link(ηDC), VSC converter efficiency (ηV SC), and overall systemefficiency (ηs). CP is prescribed for each simulation andshould not deviate substantially from expected value. It iscalculated per Eqn. (1) for each rotor individually for thedouble turbine cases. Next, ηDC is computed as the meanDC link power (PDC) divided by the mean mechanical powerPm,

ηDC =PDC

Pm

, (6)

where the overbars signify time average over two full turbinerotations. This efficiency encompasses the generator and in-verter losses. Similarly, ηV SC is calculated as

ηV SC =Pe

PDC

, (7)

in which Pe is the positive-sequence active power to the grid.Further,

ηs =Pe

12ρAU

3∞, (8)

represents the “water-to-wire” efficiency of the system. Finally,total electrical efficiency is reported as

ηe = ηDCηV SC . (9)

III. RESULTS AND DISCUSSION

Efficiency results, reported with two significant digits, forthe four simulation topologies are presented for comparisonin Table IV. Slight deviation of mean CP was noted forseveral of the simulation runs. This is due to two factors: finiteinertia of the generator model and tuning of the speed controlloops. Traces of speed and torque response with respectivecommands for coordinated intracycle angular velocity control(simulation 1) are underlaid over a short time period andpresented in Fig. 5. Actual speed tightly tracked command,while torque oscillated about the setpoint with a frequency andamplitude dependent on the rotational moment of inertia andPI gains. These variations resulted in slightly lower simulatedCP than commanded. The speed and torque inputs wereoptimized for a turbine with specific properties and adapted toa much larger power scale. In practice, controller optimizationwould be conducted on the field-scale system and would resultin speed and torque profiles somewhat unlike those used todrive this simulation.

The majority of losses in power incurred under advancedcontrol in simulations 1 through 3 were from the generator andrectification to DC. Of the two, generator losses appear moresubstantial, as the generator drive is modeled with low on-stateresistance and to minimize leakage current in the off-state.This is expected, and an acknowledged challenge of intracycleangular velocity control, which requires repeated transientaccelerations and decelerations resulting in large transientand bipolar peak currents. Additionally, the high dynamicrequirement necessitates the use of larger (i.e., more copper)generators to accommodate peak loading, and these generatorcopper losses severely impact overall system efficiency.

1.35 1.4 1.45500

1000

1500

2000

2500G

en. S

pee

d (

rpm

)US Commands (rpm/Nm)

US Actual (rpm/Nm)

DS Commands (rpm/Nm)

DS Actual (rpm/Nm)

1.35 1.4 1.45

Time (s)

-20

0

20

40

60

Gen

. T

orq

ue

(Nm

)

Fig. 5. Speed command and response for upstream/downstream turbines,simulation 1

Comparing simulations 2 and 3, utilizing two half rotors onthe same DC bus improved overall system efficiency by 1%, ora 5% increase, despite simulation 2 running with slightly lowerinput CP . This was due to substantially higher efficiency to theDC link. Though torque and current magnitudes were halvedfor this half-rotor case, copper losses in the generators werelower, as these are proportional to the square of current. Thishighlights a benefit of reducing peak current generation anddemand by splitting up power between two units. Efficiencyof the system at the VSC was identical for these two cases.

The highest performance was obtained during simulation1. Much of this is ascribed to the downstream rotor’s higherinput CP . However, the improvement in ηDC and ηV SC canbe linked to reduced generator losses and power combination.Transient current generation and demand was reduced by thetorque (current) production phase of one turbine being coor-dinated with the torque demand phase of the other, loweringresultant peak DC link current. System efficiency gain overthe constant speed case (simulation 4) is a modest 13% forcoordinated intracycle angular velocity control with two half-rotors, despite 31% greater input CP (if average of upstreamand downstream CP is considered as array efficiency). Furtherefficiency gains may be obtained by altering the objectiveof controller optimization to target ηs as opposed to CP .For example, a constraint may be added to minimize oreliminate the controller’s ability to draw power from thegrid. Economic factors such as capital cost of hardware mustbe considered when weighing the benefits of the evaluatedadvanced controllers.

Mean and extreme values of power to the grid are pre-sented in Table V. Simulation 1 generates the highest mean

TABLE IVSIMULATED SYSTEM EFFICIENCY SUMMARY

Sim. 1 Sim. 2 Sim. 3 Sim. 4US CP

∗ 0.30 0.30 0.30 0.26DS CP

∗ 0.38 - - -US CP 0.29 0.29 0.30 0.25DS CP 0.36 - - -ηDC 0.88 0.86 0.78 0.97ηV SC 0.96 0.91 0.91 0.99ηe 0.85 0.78 0.71 0.95ηs 0.27 0.22 0.21 0.24

* signifies setpoint, US/DS signify upstream/downstream.

0 0.5 1 1.5Time (s)

-20

-10

0

10

20

30

Po

wer t

o G

rid

(k

W)

Sim. 1Sim. 2Sim. 3Sim. 4

Fig. 6. Positive-sequence active power to grid

power, while halving transient spikes of generation and drawcompared to simulations 2 and 3, as shown in Fig. 6. Thisfeature is critical for the rating of associated power electronicsand peripheral hardware of the system, allowing for lower-cost components and connections. As expected, constant speedcontrol yields the most consistent product to the grid. Ad-ditionally, the table reports deviation from DC link voltagesetpoint as low in all cases, as well as acceptable maximumabsolute turbine speed tracking error (ωerr). The high peak-to-average ratio of power to and from the grid under advancedcontrol may be problematic from a power quality perspective.Research into energy storage for smoothing coupled to theDC bus has the potential to alleviate this, and is the subjectof ongoing research as a continuation of this work.

The development of this simulation parallels the develop-ment of emulation hardware to test these advanced controlschemes in a laboratory environment. Electromechanical em-ulation offers the utility of evaluating actual generators andpower electronics that are designed for or are candidates forintegration in field-deployed turbine systems [11]. Resultsfrom the simulation will guide remaining design decisionsfor sizing hardware components and generating test plansand parameters. Efficiencies obtained in emulation will becompared to those presented here in a future publication.

A subset of the power electronics comprising the collectionscheme analyzed must be co-located with turbines on theseabed for deployment lacking surface expression. At mini-mum, this include generators and their individual inverters,

TABLE VSIMULATION POWER, VOLTAGE, AND TRACKING SUMMARY

Sim. 1 Sim. 2 Sim. 3 Sim. 4Max Pe

(kW) 10.7 22.4 22.1 3.69

Min Pe

(kW) -7.58 -15.7 -15.3 -1.15

Mean Pe

(kW) 1.10 0.893 0.857 0.972

Max VDC

(V) 602.7 603.3 603.6 603.9

Min VDC

(V) 598.2 598.2 598.2 597.6

Mean VDC

(V) 600.0 600.0 600.0 600.0

Max ωerr 2.6% 2.7% 2.8% 1.5%

breakers, and associated cabling. Losses in the generatorsmanifest as heat, and at the power levels reported, impose achallenge for effective dissipation in sealed pressure housings,likely necessitating an active cooling subsystem. Subsea powercabling projects as a significant cost for a field deployment ofa turbine array. Cable runs can be minimized by additionallysubmerging the VSC inverter and controller as a DC node,allowing short runs of cable between individual turbines andthe node, and running a single three-phase AC line to ashore station. A significant disadvantage of this layout isthe requirement of submerging storage elements to interfacewith the DC bus to avoid additional losses incurred from arectification stage on shore. Additionally, failure at the VSCwould render the entire array inoperable, possibly necessitatingmultiple units in parallel. Indeed, in a review assessing the re-liability of subsea power systems for utility-scale tidal energy,Narayanaswamy and Bang-Andreasen estimate AC/AC con-verters, cooling systems, and VSCs as components most likelyto fail within five years [12]. These design considerations willbe weighed with component cost during the planning stagesof a field-deployed version of this system.

IV. CONCLUSION

A numerical simulation of power collection from multiplehydrokinetic turbines is presented. A total system efficiencygain of 13% is achieved by collecting generated power froma pair of turbines running an advanced coordinated controlscheme, compared to a single turbine with an equivalent powerrating running at constant speed. Losses in the generators arefound to be the most significant factor determining electricalconversion efficiency, and high, transient peaks of currentresult in lower generator efficiency for systems employingadvanced controllers. Losses are reduced under coordinatedcontrol, where the power generation phase of one turbine iscotemporal with the power consumption phase of a second.However, spikes of power up to ten-times greater than meanpower output are still reported for this scheme, underscoringthe challenge of balancing optimized mechanical rotor perfor-mance with obtaining reasonable total water-to-wire efficiency.

ACKNOWLEDGMENT

This work is supported by Naval Facilities Engineer-ing Command under award N00024-10-D-6318. The authorswould like to thank Benjamin Maurer for his review andsuggestions.

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