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ΕΘΝΙΚΟ ΜΔΣΟΒΙΟ ΠΟΛΤΣΔΥΝΔΙΟ
ΣΥΟΛΗ ΗΛΔΚΣΡΟΛΟΓΩΝ ΜΗΥΑΝΙΚΩΝ ΚΑΙ
ΜΗΥΑΝΙΚΩΝ ΥΠΟΛΟΓΙΣΩΝ
ΤΟΜΔΑ ΗΛΔΚΣΡΙΚΗ ΙΥΤΟ
Ανάπηςξη ηοπολογίαρ πποζαπμοζηικού ελέγσος διεπγαζιών πλοίων με
πςθμιζηέρ ζηποθών για εξοικονόμηζη ιζσύορ και καςζίμος γεννηηπιών
Π Δ Ρ Ι Λ Η Φ Η ........................................................................................................................................................ 5
A B S T R A C T ........................................................................................................................................................ 6
[J1] S. V. Giannoutsos and S. N. Manias, "A Data-Driven Process Controller for Energy-Efficient
Variable-Speed Pump Operation in the Central Cooling Water System of Marine Vessels", in IEEE
Transactions on Industrial Electronics, vol. 62, no. 1, pp. 587-598, Jan. 2015
[J2] S. V. Giannoutsos and S. N. Manias, "A Systematic Power-Quality Assessment and Harmonic Filter
Design Methodology for Variable-Frequency Drive Application in Marine Vessels", in IEEE
Transactions on Industry Applications, vol. 51, no. 2, pp. 1909-1919, March-April 2015
[J3] S. V. Giannoutsos and S. N. Manias, "Improving Engine Room Ventilation Systems: A Data-Driven
Process Controller for Energy-Efficient, Variable-Speed Fan Operation in Marine Vessels", in IEEE
Industry Applications Magazine, vol. 22, no. 6, pp. 66-81, Nov.-Dec. 2016
[C1] S. V. Giannoutsos and S. N. Manias, "Energy management and D/G fuel consumption optimization
in the power system of marine vessels through VFD-based process flow control," 2015 IEEE 15th Int.
Conference on Environment and Electrical Engineering (EEEIC), Rome, 2015, pp. 842-850.
[C2] S. V. Giannoutsos, S. Kokosis and S. N. Manias, "A gate drive circuit for Normally-On SiC JFETs
with self-protection functions against overcurrent and shoot-through fault conditions," 2015 IEEE
15th Int. Conference on Environment and Electrical Engineering (EEEIC), Rome, 2015, pp. 851-859.
[C3] S. V. Giannoutsos and S. N. Manias, "Development of an integrated energy efficiency control
system for ship power balance and diesel generator fuel consumption optimization," 2013 IEEE
Industry Applications Society Annual Meeting, Lake Buena Vista, FL, 2013, pp. 1-11
[C4] S. V. Giannoutsos and S. N. Manias, "A cascade control scheme for a grid connected Battery Energy
Storage System (BESS)," 2012 IEEE International Energy Conference and Exhibition
(ENERGYCON), Florence, 2012, pp. 469-474.
IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 62, NO. 1, JANUARY 2015 587
A Data-Driven Process Controller forEnergy-Efficient Variable-Speed PumpOperation in the Central Cooling Water
System of Marine VesselsSpyridon V. Giannoutsos, Student Member, IEEE , and Stefanos N. Manias, Fellow, IEEE
Abstract—In this paper, a data-driven process controlleris designed and implemented onboard a typical marinevessel for optimal variable-speed pump operation, leadingto the energy efficiency optimization of its central coolingwater system. To match variable flow rate requirementsdue to changes in the vessel’s operational profile with re-spect to plant limitations, real-time process measurementsare used as feedback signals to adjust the parametersand set-points of a data-driven process controller withself-tuned proportional–integral–differential control loops,which is realized through a commercial programmable logiccontroller and regulates the speed of sea water coolingpumps in order to maximize power saving potential dur-ing sea-going and cargo unloading periods. Data-drivencontrol establishes system dynamics according to processparameter variation and ensures system reliability throughparameter monitoring, regardless of the controlled plantmodel. The plant power saving potential is initially exam-ined through a practical case study, whereas experimentalresults provided after the proposed control system retrofitinstallation onboard a tanker vessel show significant powerbalance improvement and reduction of diesel generator fuelconsumption compared to existing pump throttle controlmethods, verifying that marine industry can be greatlybenefited from this energy efficiency upgrade.
Index Terms—Data-driven control, flow control, indus-trial power system control, marine vehicle power systems,power conversion, process control, pumps, variable-speeddrives.
I. INTRODUCTION
IN order to reduce fuel consumption, marine vessels nowa-days use slow steaming, operating main engine (M/E) at
lower speeds [1]. This practice leads to variable flow require-ments of auxiliary machinery associated with water cooling,
Manuscript received August 30, 2013; revised December 14, 2013and January 28, 2014; accepted March 20, 2014. Date of publicationApril 14, 2014; date of current version December 19, 2014. This workwas supported by Thenamaris Ships Management, Inc. as part of itsstrategic plan for optimization of energy management across its fleet.
S. V. Giannoutsos is with Thenamaris Ships Management, Inc.,16671 Athens, Greece, and also with the Department of Electrical andComputer Engineering, National Technical University of Athens, 15780Athens, Greece (e-mail: [email protected]; [email protected]).
S. N. Manias is with the Department of Electrical and ComputerEngineering, National Technical University of Athens, 15780 Athens,Greece (e-mail: [email protected]).
Color versions of one or more of the figures in this paper are availableonline at http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TIE.2014.2317456
ventilation, and air conditioning systems due to the need forderated capacities [2], [3]. However, for pumps used in marinevessels, flow regulation is currently achieved by throttlingthrough valve operation in their discharge side, which is energyinefficient since the power consumption of the pump motor isnot affected. Previous economic studies suggest that applyingvariable-frequency drives (VFDs) for flow regulation in pump-ing systems can lead to significant energy savings, particularlyin marine vessels [4], [5].
While energy efficiency studies demonstrate the importanceof using variable-frequency control in industrial applications inorder to improve energy efficiency [6]–[11], effective inductionmotor drive control strategies focusing on current regulationand vector control have been proposed for pumping systems[12]–[14], and several options are presented for electric ships[15]. With regard to the performance of adjustable-speed cen-trifugal pumps, the effects of voltage variation and unbalanceconditions are investigated in [16] and [17]. A hybrid estima-tion method for the centrifugal pump operational state whenfrequency converters are used is presented in [18], using flowrate and head estimates from the process. However, the men-tioned model-based control approaches are based on sufficientquantitative knowledge being available from the process and arelimited to the frequency converter or motor itself, without con-sidering the effects on plant dynamics and constraints, whichare important in complex industrial processes that combinehydraulic, mechanical, and electrical systems, as it is often thecase, particularly in marine vessels.
Data processing, on the other hand, can be used for faultdetection and diagnosis in such complex industrial processes,as proposed in [19]. To this end, a real-time data-driven im-plementation of fault tolerant control system, which uses anobserver/residual generator of fault diagnosis vectors, is pre-sented in [20] and [21]. To achieve robust control that is modelfree, a data-based state feedback control method using a statesignal sampling technique is proposed in [22], whereas a dy-namic linearization-based data-driven controller is presented in[23]. In [24], the tuning of a fixed-order data-driven controllerfor linear time-invariant plants is achieved through the use ofinput–output samples.
In this paper, a data-driven process control method is in-troduced to optimize power consumption of adjustable-speed-driven centrifugal sea water (S.W.) cooling pumps in thevessel’s central cooling water system in relation to process
588 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 62, NO. 1, JANUARY 2015
variable flow rate requirements. The controller parameter ad-justment is based on measured process input and output tem-perature and pressure data, which are used as feedback signalsin self-tuning proportional–integral–differential (PID) controlloops, whereas the outputs are bounded accordingly in orderto comply with varying process constraints, improving systemstability. The controller dynamically changes the pump speedset-point in relation to the vessel’s operating profile, corre-sponding to sea-going and cargo handling periods, respectively,in order to achieve plant energy efficiency through power bal-ance optimization with respect to real-time M/E and auxiliarymachinery water flow requirements.
To examine the plant power saving potential and deter-mine the critical process parameters, case study results arepresented from the modeling of the vessel’s central coolingwater system according to derated M/E and auxiliary machineryflow requirements, which are related to slow steaming prac-tice. The proposed pump control scheme is realized througha commercial programmable logic controller (PLC), whereasa human–machine interface (HMI) is implemented for plantparameter monitoring. Experimental results derived from sys-tem retrofit installation onboard a typical tanker vessel showthat the controller dynamics are well established, leading tosignificant energy savings as well as to diesel generator (D/G)fuel consumption and emission factor reduction in comparisonto existing pump throttling control.
II. OPTIMAL FLOW REGULATION FOR
PUMPS AT SLOW STEAMING
A fluid flow system is characterized by a system curve, wherethe system head h (in ft) is a function of the static head dhand where head losses hl depend on the flow velocity q (inft/min) on acceleration due to gravity g (in ft/min2) and on theconstant k describing the friction caused by the piping system,as expressed in
h = dh+ hl = (h2 − h1) +kq
2g
2
. (1)
To service the fluid flow system, the operating point of thedesigned pump is the intersection of the system curve andthe pump performance curve, with associated motor powerconsumption PkW (in kW), as expressed in
PkW = 0.746×(
γ × v × h
33 000× η × ηe
)(2)
where γ is the specific weight of the fluid (in lb/ft3); v is thevolumetric flow rate (in ft3/min); and η and ηe are the overallpump efficiency and motor efficiency factors, respectively.
Centrifugal pumps onboard vessels are always designed toprovide the required flow rate for M/E operation at the maxi-mum capacity ratio (MCR); therefore, the operating point of thesystem is originally A (hA, vA), where the pump motor is work-ing at nominal speed ωMCR in order to service the system de-scribed by curve sa, as shown in Fig. 1. Through the closing op-eration of valves in the piping system, the system curve movesupward to sa′ ; thus, the operating point moves along the samepump curve to A′ (hA′ , vA′), achieving flow reduction at higher
Fig. 1. Dynamic power saving margin according to pump performanceand system curve variation at different flow rate requirements.
impeller pressure, which is energy inefficient, as pump powerconsumption remains almost nominal, i.e., PMCR, as before.
Due to the vessel’s slow steaming operating practice, M/Eoften operates at lower speeds, resulting in less flow raterequirements from the process. In that case, variable-frequencycontrol can be used to adjust the speed of the pump motor to alower value ωN by moving the whole pump curve downward sothat the operating point moves from A to B (hB , vB) or fromA′ to B′ (hB′ , vB′) accordingly, resulting in significant powersavings as highlighted in the respective areas in Fig. 1, since thepump power consumption PN is reduced according to affinitylaws, as described in
ωN
ωMCR=
νsw_in,N
νsw_in,MCR= 3
√PN
PMCR. (3)
However, since the pump operation is associated with aspecific process whose parameters are varying according to thevessel’s operating profile, M/E speed, auxiliary machinery con-dition, and physical constraints, adjustment of ωN simply basedon process modeling is not adequate without the risk of jeopar-dizing integrity and reliability. To this end, the objective of theintroduced data-driven controller is to self-tune via real-timeprocess measurements in order to operate the pump at an op-timal speed, so that maximum power savings are achieved andthe minimum speed limit ωmin is adapted in compliance withthe minimum required flow limits vmin, which vary accordingto the needs of M/E and auxiliary machinery at each time.The design and implementation of the data-driven controlleris performed for energy efficiency optimization of the vessel’scentral cooling water system. Therefore, the process model isobtained to examine the process constraints and investigate therequirements through a case study for the vessel where thecontroller will be implemented as a retrofit installation.
III. MODELING OF THE VESSEL’S CENTRAL COOLING
WATER SYSTEM CAPACITIES FOR M/ESLOW STEAMING OPERATION
Since the process parameters considerably vary in relationto the M/E operating point, in order to develop the modeland assess the capacities of the vessel’s central cooling watersystem, the M/E speed and load variations in the slow steamingcondition have to be considered. For this reason, the engine
GIANNOUTSOS AND MANIAS: DATA-DRIVEN PROCESS CONTROLLER FOR VARIABLE-SPEED PUMP OPERATION 589
Fig. 2. M/E load diagrams for the ship’s slow steaming operating point.
layout and load diagrams for the vessel under investigationare presented in Fig. 2, along with the comparison betweenthe corresponding MCR, continuous nominal (SP), and slowsteaming operating points (N).
Since a fixed pitch propeller is used, engine power PM/E
varies with propeller speed n according to
PM/E = c× n3, where c = constant. (4)
It is also known that the effective power of a marine dieselengine depends on mean effective pressure pe, as in
PM/E = c× pe × n = c′ × n, for constant pe. (5)
For the vessel under investigation, plotting the percentagesof M/E power for a range of operating speeds according to(4) and (5) in logarithmic scale along with the two limitingconstant pe lines, i.e., L1−L3 and L2−L4, and the two constantspeed lines, i.e., L1−L2 and L3−L4, results in engine layoutand load diagrams, as shown in Fig. 2. While the MCR pointrefers to nominal speed (nMCR = 100%), the engine contin-uous operating point SP is determined on the heavy runningand fouled hull propeller curve after incorporating a 5% heavyrunning margin and a 15% rough sea margin. However, around90% of the vessel’s sea-going period, slow steaming is used toreduce M/E fuel consumption; thus, the operating point movesalong the heavy running curve to point N, resulting in 84.4% ofnominal speed. The M/E MCR point is originally used duringthe design stage of the vessel’s central water cooling system,which is modeled in Fig. 3, to dimension central cooler capacityQcent,MCR (in kW).
The central cooler in service is actually a heat exchanger,compensating for the MCR capacities required by major en-gine room consumers for heat dissipation purposes, involvingcapacities of M/E scavenge air cooler Qair,MCR, which isconnected at the F.W. side in parallel with jacket water coolerQjw,MCR and lubricating oil cooler Qlub,MCR, respectively. Onthe S.W. side of the central cooler, a S.W. cooling pump isdesigned to operate in constant speed and supply water withflow vSW_in,MCR, considering a S.W. temperature of 32 ◦C.The S.W. flow is also used during the sea-going period forcondensating high-temperature M/E exhaust gases through anatmospheric condenser, whereas in the cargo discharging pe-
Fig. 3. Simplified model of the vessel’s central water cooling system.
riod, a vacuum condenser uses S.W. to expand steam to lowerpressure in order to drive the steam turbines. The temperatureon the heat exchanger F.W. outlet side is currently regulatedto be 34 ◦C through a thermostatic three-way valve, whichdetermines if F.W. flow should bypass the central coolers, basedon F.W. temperature measurement.
However, for usual M/E operation at slow steaming point N,the heat dissipation requirements of the mentioned consumersare significantly lower; thus, the corresponding heat dissipationreduction factors can be calculated according to
where nN% and PN% are the percentages of M/E speed andload, respectively, at point N [25].
The reduced consumer required capacities can be calculated(in kW) through the following equations for derated M/E oper-ation at point N:
Qair,N =Qair,MCR ×Qair,N% (9)
Qjw,N =Qjw,MCR ×Qjw,N% (10)
Qlub,N =Qlub,MCR ×Qlub,N%. (11)
The variation of heat dissipation reduction factors at differentoperating points, including selected point N, is shown in Fig. 4.
The required central cooler capacity at slow steaming M/Eoperating point N, i.e., Qcent,N , can be expressed as
Qcent,N = Qair,N +Qjw,N +Qlub,N . (12)
Therefore, the corresponding required water flow from the S.W.pump at operating point N, i.e., vSW_in,N , is proportional to therequired central cooler capacity, as shown in
νSW_in,N = νSW_in,MCR × Qcent,N
Qcent,MCR(13)
where vSW_in,MCR and Qcent,MCR are the designed S.W. flowand central cooler capacities, respectively, at MCR. From (13),it is shown that the required water flow from the S.W. pumpsignificantly varies with M/E load and speed.
590 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 62, NO. 1, JANUARY 2015
Fig. 4. Variation of heat dissipation reduction factors for major E/R consumers in the vessel’s central cooling water system at different M/E operatingpoints. (a) Scavenge air cooler heat dissipation reduction factor at slow steaming point N, Qair,N%. (b) Jacket water cooler heat dissipation reductionfactor at slow steaming point N, Qjw,N%. (c) Lubricating oil cooler heat dissipation reduction factor at slow steaming point N, Qlub,N%.
TABLE ICENTRAL COOLING WATER SYSTEM DESIGNED CAPACITIES
TABLE IIREQUIRED SYSTEM CAPACITIES AT OPERATING POINT N
For the tanker vessel under investigation, Table I showsthe central cooling water system designed capacities for M/Eoperation at MCR, whereas the required system capacities
for slow steaming operation at point N are calculated andpresented in Table II, based on the developed model. As shownin Table II, although the flow rate required to be provided by
GIANNOUTSOS AND MANIAS: DATA-DRIVEN PROCESS CONTROLLER FOR VARIABLE-SPEED PUMP OPERATION 591
Fig. 5. Proposed data-driven controller interface with the process.
the S.W. cooling pump at slow steaming operation at pointN is reduced to vsw_in,N = 57.4% compared to MCR, thepump motor always consumes its nominal power of PSW =75 kW. If variable-frequency control was simply used to adjustthe pump speed manually in order to reduce its flow to thatpoint, theoretically, its power consumption could be reducedto PSW = 14.2 kW according to (3). However, this practicealone may not satisfy several important process parameters andcritical constraints, which vary among sea-going, anchorage,loading, and unloading vessel operating conditions.
The changes in S.W. temperature, variations in S.W. pres-sure provided to the central coolers and to the atmosphericcondenser for proper operation, minimum F.W. temperatureand pressure limits for M/E and D/G jacket cooling, centralcooler performance and piping system condition, as well asvacuum (C.O.P.T.) condenser outlet temperature variation mustall be taken into account by the controller, which will providethe pump speed limit. The proposed data-driven controllerminimizes pump power consumption while satisfying the abovedynamically changing parameters without relying on the modelof the process. Override functions in case of fault detection arealso integrated into the controller.
IV. DESIGN, CONTROL LOGIC, AND TUNING OF THE
PROPOSED DATA-DRIVEN PROCESS CONTROLLER
The proposed data-driven controller interface with the pro-cess is presented in Fig. 5 using sets of real-time measurementsto regulate parameter set-points and set the pump motor speedaccordingly.
During the vessel’s sea-going, anchorage, and loading peri-ods, temperature measurements are obtained from the F.W. sideof the central coolers after the thermostatic three-way valve,i.e., TF.W._out (∼◦ C), and are provided as a control input tothe proposed process controller. This variable is used to controland maintain a desired F.W. system temperature reference,i.e., T ∗
F.W._out, which has a direct impact on the cooling ofM/E, D/G, and auxiliary machinery. The three-way v/v posi-tion, i.e., v/vpos (∼ %), is monitored and adjusted based onthe T ∗
F.W._out set-point in order to achieve a better system
dynamic response. In addition, the level of S.W. pressure onthe discharge side of the S.W. pump, i.e., PS.W._out (∼ bar), isalso provided as an input to the controller in order to maintaina desired pressure reference, i.e., P ∗
S.W._out, considering themeasurement of S.W. inlet temperature in the suction side ofS.W. pumps, i.e., TS.W._in (∼◦ C), in order to avoid creationof vacuum condition in the atmospheric condenser and complywith the required pressure limits in the S.W. and F.W. sys-tems. A minimum speed limit ωmin_sea is introduced based onmeasurements and physical constraints to avoid fouling in thepiping system and central coolers.
In the cargo unloading period, since steam-driven cargopumps are in service, a C.O.P.T. condenser is installed inthe rear of the turbine, using S.W. to decrease back pressureand retrieve steam for the purpose of recycling it into boilingwater in order to increase turbine efficiency. During this mode,the controller objective is to maintain a constant temperaturedifference T ∗
S.W._diff , between the outlet and inlet sides ofthe condenser, according to measurements of the condenserS.W. outlet temperature TS.W._out (∼◦ C) and the S.W. inlettemperature TS.W._in (∼◦ C). Since M/E is not in use, coolingrequirements are reduced, but at the same time, a different speedlimit ωmin_cargo has to be introduced based on S.W. pressure,since water flow must service both the cooling system of D/G’sand auxiliary machinery and ensure C.O.P.T. condenser properoperation.
In Fig. 6, the layout of the proposed data-driven processcontrol algorithm and the related pump variable-frequency con-trol topology are presented with reference to the vessel underinvestigation. The applied variable-frequency control topologyconsists of a standard six-pulse diode rectifier and a two-levelthree-phase voltage source inverter (VSI). It includes 3% dcand ac chokes to maintain total voltage harmonic distortion(THDv%) within the limits specified by marine classificationsocieties.
During sea-going, anchorage, and cargo loading periods,the desired T ∗
F.W_out and P ∗S.W._out set-points are maintained
through two independent PID control loops, which are usedto regulate motor speed references ω∗
T_F.W. and ω∗P_S.W.,
respectively, according to
ω∗T_F.W. =K∗
P,T × eT_F.W.(t) +K∗I,T ×
t∫0
eT_F.W.(t) dt
+K∗D,T × d
dteT_F.W.(t) (14)
ω∗P_S.W. =K∗
P,P × eP_S.W.(t) +K∗I,P ×
t∫0
eP_S.W.(t) dt
+K∗D,P × d
dteP_S.W.(t) (15)
where eT_F.W.(t) and eP_S.W.(t) are expressed by
eT_F.W.(t) =T ∗F.W._out − TF.W._out(t) (16)
eP_S.W.(t) =P ∗S.W._out − PS.W._out(t). (17)
To comply with dynamic system changes and process con-straints, the output speed command signal from the above PID
592 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 62, NO. 1, JANUARY 2015
Fig. 6. Design of the proposed data-driven process control topology.
control loops, i.e., ω∗T_P , is the maximum value between the
two references, whereas the final speed command provided tothe pump motor during this period, i.e., ω∗
sea, is bounded byωmin_sea, as in
ω∗T_P = max
{ω∗T_F.W., ω
∗P_S.W.
}, ωmin_sea ≤ ω∗
sea ≤ 95%.(18)
Data-driven control is used to automatically adjust T ∗F.W._out
and P ∗S.W._out set-points according to the TS.W._in range, as
well as dynamically select ωmin_sea so that M/E and D/G jacketcooling water temperature and pressure remain below 85 ◦C andabove 3.6 and 3.4 bars, respectively, whereas vacuum conditionand temperature rise in atmospheric condenser are avoided bykeeping its outlet S.W. pressure above 0.5 bar. The controllerparameters, set-points, and minimum limits at the sea-goingperiod are defined in Table III.
To improve control system response and further reduceωmin_sea imposed by the controller, the temperature setting ofthe thermostatic three-way v/v, i.e., T ∗
3-way, is adjusted 2 ◦Clower in comparison to T ∗
F.W._out, for the three-way v/v to openand allow more water to flow from central coolers, stabilizingthe plant at pump reduced speeds, as in
T ∗3-way = T ∗
F.W._out − T ∗sett(t) = T ∗
F.W._out − 2 ◦C. (19)
Based on T ∗3-way defined above, the v/vpos% is also measured.
During the cargo unloading period, the desired differentialtemperature set-point T ∗
S.W._diff is maintained through a PID
TABLE IIIDEFINITION OF CONTROLLER PARAMETERS
AT SEA-GOING CONDITION
control loop, which is used to regulate speed reference ω∗T_diff ,
To comply with system changes and process constraints, thefinal speed reference in this mode, i.e., ω∗
cargo, is bounded byminimum and maximum limits, as presented in
ωmin_cargo ≤ ω∗cargo ≤ 95%. (22)
Data-driven control is used to automatically adjust theT ∗S.W._diff set-point according to the PS.W._out range, as well
as dynamically select ωmin_cargo in order to compensate forthe need to keep S.W. flowing both for D/G and auxiliarymachinery cooling and for vacuum condenser operation. In casepressure drops below a minimum limit of 0.6 bar, which is theexisting S.W. pressure switch set-point, the variable-frequencycontrolled pump motor runs at a maximum of 95% of ωnom
to avoid start of the standby pump. The controller dynamicparameters, set-points, and minimum limits during the cargohandling period are defined and shown in Table IV.
Regarding recently proposed PID controller tuning methods,a generalized transfer function process model consisting ofanalytical expressions of process parameters was used in [26],whereas in [27], a fuzzy saturation-based tuning method usingsensitivity functions was proposed. Although the mentionedmethods demonstrate effectiveness in response, they require adetailed plant model, increasing computational complexity. Inthis kind of application where the plant is subject to changesin its layout, operation, and parameters during the operatingprofile, with slowly changing states such as water temperatureand pressure, the gains of PID controllers incorporated in the
GIANNOUTSOS AND MANIAS: DATA-DRIVEN PROCESS CONTROLLER FOR VARIABLE-SPEED PUMP OPERATION 593
TABLE IVDEFINITION OF CONTROLLER PARAMETERS AT CARGO TANK OPERATION
control scheme are designed to be self-tuned according tovalues obtained through real-time process measurements.
It was shown that ωmin was dynamically defined accordingto plant requirements and process constraints for each operatingperiod. In this case, the ωmin/ωmax ratio can be used to tuneK∗
P . Lower ωmin results in less water flow rate and pressure inthe system, requiring an increase in rise time and decrease inovershoot by lowering K∗
p in order to protect the process fromsudden peaks that may appear in S.W. pressure from a changein the speed set-point and vice versa. Moreover, the three-way thermostatic valve position v/vpos set by the automaticallydefined T ∗
F.W._out set-point is used to regulate K∗D. In case it
opens to the central cooler side, v/vpos% is reduced, providingthe opportunity to increase K∗
D in order to decrease correspond-ing overshoot and settling time of the controller response. Thementioned self-tuning data-driven approach is expressed by
K∗P =(0.60×Ku)×
ωmin
ωmax(23)
K∗I =2× K∗
P
Tu(24)
K∗D =
(K∗
P × Tu
8
)× 1
10× (ν/νpos%/100)(25)
where Ku is the ultimate gain that the output oscillates atconstant amplitude with period Tu. For the initial tuning, thepump operates at ωmax = 95%, whereas v/vpos% = 100%.
A built-in fault detection mode is implemented in thecontroller, where in case there is a sensor, control unit, orcommunication failure, a system override function is activated,and the speed reference is ω∗
fault = 95%. Safety checks areincluded, so that even if in this condition PS.W._out remainsbelow 0.6 bar, VFD fails, or TS.W._diff is greater than 10 ◦C,the standby S.W. pump is activated. To illustrate the differentcontrol algorithms embedded in the controller, the first cycleof the control system evolution is presented in Fig. 7 as asequential function chart (SFC).
V. ARRANGEMENT AND OPERATION OF THE PROPOSED
DATA-DRIVEN PROCESS CONTROLLER
The proposed process control topology for the optimizationof the central cooling water system is built for a typical Afra-max size tanker vessel in order to control the speed of one
Fig. 7. SFC showing the first cycle of the proposed data-driven processcontrol system evolution.
Fig. 8. Arrangement and interface of the proposed control scheme.
75-kW 440-V 60-Hz S.W. cooling pump motor, according tothe proposed control algorithm. The arrangement of the processcontrol unit as well as its interface with sensors, VFD, and HMIunits are shown in Fig. 8.
The control unit includes a Modicon M340 commercial PLCautomation platform unit with its associated analog and digitalinput–output racks and communication modules, interfacingwith an ATV61HD90N4 VFD unit. Process monitoring of theworking plant is performed in a HMISTU855 graphic displayunit, using Vijeo Designer HMI configuration software. Analoginput signals are received by 4–20 mA temperature and pressuretransmitters, which are used to collect the required process dataand provide them as real-time feedback signals to the controller.The range of PT100 temperature sensors is 0–100 ◦C, whereasthe pressure sensor range that records the S.W. pump outlet
594 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 62, NO. 1, JANUARY 2015
Fig. 9. Onboard implementation of the proposed process control topol-ogy. (a) Hardware arrangement of the proposed data-driven processcontroller. (b) Arrangement of the applied pump variable-frequencycontrol topology.
Fig. 10. Process monitoring functions of the controller HMI screen invarious modes of operation. (a) Vessel sea-going/port mode operationand process monitoring. (b) Vessel cargo unloading mode operation andprocess monitoring.
pressure is 0–6.0 bars, to compensate for the 4.0-bar nominalpump discharge pressure value. A pressure transmitter with arange of 0–1.5 bars is used to record three-way v/v positionpneumatically from the existing actuating signal. An analog4–20 mA controller output signal is used to provide the speedreference to the VFD embedded microcontroller module. Mod-bus over TCP/IP is used to synchronize VFD operation with thecontrol unit. The 24-Vdc digital output signals from the controlunit are used to control the start/stop and auxiliary functionsof the VFD, whereas the digital input signals from VFD informthe controller about the system status, including overload, phasefailure, temperature rising, and VFD fail conditions.
The hardware configuration of the process control unit in-stalled onboard the vessel is shown in Fig. 9(a), whereas thearrangement of the VFD topology, which was built inside apanel in compliance with classification society regulations, ispresented in Fig. 9(b). HMI process monitoring functions areshown in Fig. 10 for various operating modes. Through theHMI, it is noted that the user can monitor and adjust importantprocess parameters as well as apply different process controlalgorithms by selecting between sea-going/port and cargo tankoperation modes at every time.
In Fig. 10(a), it is shown that in the sea-going mode, whilethe vessel operates at near 70% of the nominal M/E load(∼10 000 kW), the controller operates the pump at 72.0% ofits nominal speed, resulting in 28 kW power consumptionor 47 kW power savings compared to the previous throttlingoperating practice, whereas TF.W_out is kept below 34 ◦C,PS.W_out is maintained at 1.19 bars, and TS.W._in is 20.4 ◦C.It is noted that since the C.O.P.T. condenser is not in service at
Fig. 11. Dynamics of the sea-going/port mode process control.(a) Central cooler F.W. outlet temperature PID control (K∗
Fig. 12. Dynamics of the cargo unloading mode process control.(a) Central cooler F.W. outlet temperature PID control (ignored).(b) Differential temperature PID control (K∗
P,diff = 24.5, K∗I,diff =
2.5, K∗D,diff = 1.6).
that point, S.W. inlet and S.W. outlet temperatures are the same.The v/vpos% is adjusted to 83.0% at the central cooler bypassto achieve higher F.W. pressure since TF.W._out is below 34 ◦C.
When the vessel is subject to cargo discharging operation,unloading mode control takes over, as shown in Fig. 10(b).In this mode, the pump motor operates at 80% of its nominalspeed, leading to 38 kW power consumption or 37 kW powersaving from the vessel’s D/G load. It is shown that the controllerignores the set-points of the previous mode and maintains aT ∗S.W._diff of 3.39 ◦C, due to use of the cargo condenser. It
is noted that PS.W._out has dropped to 0.73 bar because theS.W. pump now services both the central cooling water systemand the cargo condenser. This condition imposes 80% as thelower speed limit in this mode, due to the minimum S.W. outletpressure requirement of 0.6 bar.
VI. EXPERIMENTAL RESULTS FROM THE PROPOSED
DATA-DRIVEN CONTROLLER RETROFIT
INSTALLATION IN A TANKER VESSEL
A. Controller Dynamics in Various VesselOperating Conditions
The performance of the proposed data-driven controller withregard to process parameter variations is examined and vali-dated through evaluation of its dynamic responses, as shown inFigs. 11 and 12 for sea-going/port and cargo unloading modesof operation. The controller override functions during the faultdetection mode are presented in Fig. 13.
GIANNOUTSOS AND MANIAS: DATA-DRIVEN PROCESS CONTROLLER FOR VARIABLE-SPEED PUMP OPERATION 595
Fig. 13. Process controller override function dynamics (ω∗ref = 95%).
(a) Sea-going mode control override due to Modbus communicationfailure. (b) Cargo unloading mode control override due to temperaturesensor failure.
In Fig. 11(a) and (b), for the sea-going mode, it is shownthat since TS.W._in is 20.4 ◦C, ωmin_sea is adjusted to 65%,whereas T ∗
F.W._out and P ∗S.W._out are automatically adjusted
to 34 ◦C and 1.2 bars, respectively, according to Table III.Therefore, it is shown that during slow steaming M/E operationat nN = 86−88 r/min or PN = 7854−8568 kW load, the F.W.temperature PID control loop is tuned to provide a speed set-point ω∗
T_F.W., which is equal to 70% of ωnom. At the sametime, the S.W. pressure PID control loop provides a differentspeed set-point of ω∗
P_S.W., which is equal to 72.2% of ωnom,to comply with the requirements mentioned in Table III. There-fore, according to (18), the final speed reference provided tothe pump, i.e., ω∗
sea, is the maximum of the two set-points, i.e.,72.2% of ωnom. It is also shown that the three-way v/v position,i.e., v/vpos%, is dynamically changing according to the pumpoperating speed, assisting to the controller self-tuning, whoseresponse displays minimum overshoot and rise time, as well asexcellent steady-state characteristics.
In Fig. 12(a), for the cargo handling operation mode, M/E isnot working, and the C.O.P.T. condenser is in service since thesteam turbines are in use; therefore, F.W. temperature and S.W.pressure PID control loop outputs are ignored. On the otherhand, in Fig. 12(b), it is shown that since PS.W._out is equalto 0.73 bar, ωmin_cargo is adjusted to 80%, whereas T ∗
S.W._diffis automatically adjusted to 3.4 ◦C according to Table IV.Since one D/G and auxiliary machinery are working, the dif-ferential temperature PID control loop is tuned to provide aspeed set-point ω∗
T_diff equal to 86.2%, corresponding to ω∗cargo
as in (22).As shown in Fig. 13(a) and (b), in case of sensor or commu-
nication failure, the fault detection function is activated, and thecontrol system ignores all set-points in order to drive the pumpmotor at ω∗
fault, equal to 95% of ωnom to preserve the reliabilityof the system. It also noted that the minimum speed limit of thecontroller must be always above 50% of ωnom to protect themotor from overheating.
B. Power Balance Optimization With theProposed Controller
Using the previous throttling practice, the pump motor wasoperating constantly at fixed speed with maximum powerconsumption, since the Star/Delta starting method was used.With the proposed data-driven controller used to provide speed
Fig. 14. Inverter output and grid-side voltage/current waveforms forpump motor operation at 45Hz. (a) VSI-side voltage vm,LL and currentim,a waveforms. (b) Grid-side voltage vi,LL and current ii,a waveforms.
reference to the drive, the VSI generates a reference voltageand current waveform to control the speed of the motor. Ata typical pump motor operation at 75% of ωnom, or 45 Hz,the VSI-side voltage vm,LL and current im,a waveforms arepresented in Fig. 14(a), as measured in the motor stator. Thecorresponding grid-side voltage vi,LL and current ii,a wave-forms are presented in Fig. 14(b). The power quality at thepoint of common coupling (PCC) remains in high level, due toharmonic filtering considerations made during the design stage.
The S.W. pump motor active power consumption variationwith the use of the proposed data-driven controller is measuredand presented in Fig. 15 for a typical vessel eight-day voyagein comparison with its original power consumption before theretrofit installation. It is shown that during the vessel loadingperiod, where the heat dissipation requirements of engine roomconsumers are minimum since M/E is not working and only oneD/G is in use, the pump power consumption PSW is equal to23 kW, which approaches the minimum PSW corresponding tothe minimum allowed speed reference of 65%, as established bythe controller according to Table III. This leads to 52 kW powersaving over throttling control. During the sea-going period,controller performance was tested at the MCR point, where theM/E load PN is equal to 95%, at the nominal M/E operatingpoint (SP), where PN is equal to 85%, and at the usual slowsteaming condition, where PN is equal to 55%–60%. It isshown that the proposed control scheme dynamics are wellestablished during the vessel’s operating profile changes. Inparticular, at MCR, the controller operates the S.W. pump at95%, close to its nominal capacity to compensate for increasedM/E heat dissipation requirements. For vessel operation at theM/E nominal service point, it is shown that the pump operatesat 90% of ωnom, leading to PSW equal to 55 kW or to powersaving equal to 20 kW. In the typical slow steaming condition,however, the advantage of the proposed approach is moreevident, since PSW is just 25 kW, as compared to the pumpnominal power consumption of 75 kW, meaning that the con-troller operates the pump at 70% of ωnom, to achieve the waterflow required by engine room consumers, which is very closeto the minimum water flow corresponding to ωmin (65%) asdefined in Table III. This scale of power saving highlightsthe effectiveness of the controller at this period. During cargooperation, the pump operates at 92%, resulting in PSW equalto 58 kW or power saving of 17 kW, in order to maintain thewater flow required by central cooling system consumers and
596 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 62, NO. 1, JANUARY 2015
Fig. 15. Power saving from S.W. pump motor variable-frequency operation with the proposed data-driven control for a typical voyage.
Fig. 16. Power balance optimization through the normalization of voltage spikes at PCC during transients using the proposed process controlmethod.
for C.O.P.T. condenser operation, keeping T ∗S.W._diff equal to
3.4 ◦C and PS.W._out equal to 0.9 bar as well.The operation of the realized variable-frequency topology
associated with the proposed controller leads to significantpower demand normalization during transients and, particu-larly, during motor start-up, as shown in Fig. 16. In comparisonto the original pump throttling control and Star/Delta motorstarting mode, with the proposed data-driven VFD control, thevoltage spikes and sags that appeared at PCC during motorstart-up are greatly minimized. The gradual starting currentand improved transients lead to load leveling, since there is noneed to start an additional D/G to compensate for the initial300-kW demand that was previously present at start-up, leadingto D/G operation with improved specific fuel oil consumption(SFOCD/G) (in g/kWh).
C. Ship’s Energy Saving Optimization FromProposed Control
The contribution of the proposed control scheme to energyefficiency optimization of the vessel’s central cooling watersystem is discussed in this section by examining the impacton D/G power and fuel consumption as well as on emissionfactors (CO2, SO2, and NOx) over a typical year of operationthrough a comparative study. Considering that heavy fuel oil(HFO) is used as the standard D/G fuel, the relevant D/G fueloil consumption (FOCD/G) induced from S.W. pump powerconsumption is based on SFOCD/G and running time (RT; in h),as shown in
FOCD/G_HFO = PS.W. × RT × SFOCD/G. (26)
GIANNOUTSOS AND MANIAS: DATA-DRIVEN PROCESS CONTROLLER FOR VARIABLE-SPEED PUMP OPERATION 597
Fig. 17. Ship’s annual savings from the proposed data-drivencontroller.
For the D/Gs used onboard, a SFOCD/G of 192.3 g/kWh isconsidered if one D/G is running alone at 70% of its capacity,whereas a SFOCD/G of 210 g/kWh is considered if two D/G’sare running in parallel at 40% of their capacities.
The corresponding D/G emission factors due to fuel con-sumption are calculated based on the following equations:
where CFO is the conversion coefficient of fuel oil and carbonemission for marine heavy-duty diesel engines, and S% is thesulfur content included in fuel oil [28].
The comparative study is based on the operating profile of thespecific vessel, where experimental results were obtained fromthe proposed controller application. The comparative resultsover the pump throttling control method for a year of operationare presented in Fig. 17, including vessel normally sea-going,port in/out, anchorage/loading, and unloading periods. The en-ergy savings (in kWh) are calculated based on the difference inS.W. pump power consumption between the throttling methodand the proposed data-driven control for the specific RT ateach period. The D/G fuel savings in tons of HFO are thencalculated based on (26), whereas the calculation regarding thereduction in emission factors is based on (27)–(29), consideringa 2% sulfur percentage in HFO. Considering that the averageHFO price is $650/ton, the annual economic benefits from theproposed control system application can be derived.
It is shown that total annual benefits include 78 tons/yearreduction in HFO consumption, which means $50 700 eco-nomic benefits only from reduction of fuel costs. Consideringthat CO2, SO2, and NOx emissions are also reduced by 247.3,3.12, and 5.7 tons/year, respectively, the benefits for the marineindustry are clearly significant.
VII. CONCLUSION
In this paper, a data-driven process controller was designedfor energy-efficient variable-speed pump operation in the cen-tral cooling water system of marine vessels. The controllerparameter adjustment was purely based on real-time process
measurements, which were used as feedback signals to self-tune the controller and verify its compliance with processconstraints. The controller was used to dynamically change thespeed of the S.W. cooling pump motor to obtain the optimalpump performance curve, according to variable consumer wa-ter flow requirements and vessel’s operating profile in orderto increase energy efficiency. The use of data-driven controlcontributed to improve system dynamics and fault diagnosticsfor the varying system requirements regardless of the modelof the controlled plant. The proposed controller was realizedusing a commercial PLC, whereas an HMI was developed tomonitor process parameters, ensuring safety and reliability. Atheoretical case study was initially performed to examine thepower saving potential from the process. Experimental resultsprovided from the proposed control system retrofit installationonboard a typical tanker vessel validated that system dynamicresponses were well established, achieving significant energyand fuel savings, approaching 370 MWh and 80 tons HFO re-duction in consumption per year, respectively. Comparison re-sults over the throttling pump flow control method showed thatnot only power balance was improved, thus leading to signifi-cant economic and operational benefits, but also emission factorreduction was achieved, verifying that the marine industry canbe greatly benefited by the proposed energy efficiency upgrade,which can be adapted for use in other processes as well.
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[13] D. G. Holmes, B. P. McGrath, and S. G Parker, “Current regulationstrategies for vector-controlled induction motor drives,” IEEE Trans. Ind.Electron., vol. 59, no. 10, pp. 3680–3689, Oct. 2012.
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Spyridon V. Giannoutsos (S’12) was bornin Athens, Greece, in 1988. He received theDiploma degree in electrical and computer en-gineering in 2010 from the National TechnicalUniversity of Athens (NTUA), Athens, where heis currently working toward the Ph.D. degree inmarine electrical engineering, power electron-ics, and process control.
Since 2011, he has been a Research Asso-ciate with the Electrical Machines and PowerElectronics Laboratory, NTUA. In 2012, he
joined Thenamaris Ships Management, Inc, Athens, where he is holdingthe position of Fleet Support Electrical Engineer, focusing on designand implementation of energy and fuel saving solutions onboard. Hisresearch interests are in the fields of process control, power electronics,motor drive systems, and power quality.
Mr. Giannoutsos is a member of the IEEE Industrial ElectronicsSociety and a Registered Engineer in the Technical Chamber of Greece.
Stefanos N. Manias (M’85–SM’92–F’05) re-ceived the B.Eng., M.Eng., and Ph.D. de-grees from Concordia University, Montreal, QC,Canada, in 1975, 1980, and 1984, respectively,all in electrical engineering.
In 1975, he joined the Canadian BroadcastingCorporation (CBC), where he designed radioand television automation systems. From 1979to 1981, he was with Northern Telecom Canada,where he was responsible for the design ofswitching mode power supplies. In 1989, he
joined the Department of Electrical and Computer Engineering, NationalTechnical University of Athens, Athens, Greece, where he is currently aFull Professor teaching and conducting research in the area of powerelectronics and motor drive systems. He has authored more than 80IEEE and IEE publications in the areas of power electronics and motordrive systems.
Dr. Manias is the Chapter Chairman and the Founder of the GreeceSection of the IEEE Joint Industry Applications Society, Power Electron-ics Society, and Industrial Electronics Society and a member of the IEEEMotor Drives Committee. He is a Registered Professional Engineer inCanada and Europe.
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 51, NO. 2, MARCH/APRIL 2015 1909
A Systematic Power-Quality Assessment andHarmonic Filter Design Methodology for
Spyridon V. Giannoutsos, Student Member, IEEE, and Stefanos N. Manias, Fellow, IEEE
Abstract—In maritime industry, high fuel costs encourage use ofvariable-frequency drives (VFDs) for energy-saving applications.However, introduction of such nonlinear loads in the vessel’spower distribution network induces harmonics, which can lead topotential risks if are not predicted and controlled. In this paper, asystematic power-quality assessment and monitoring methodologyis proposed to calculate VFD contribution to voltage distortionat the point of common coupling (PCC), considering the sourceshort-circuit capacity and the existing vessel’s power system har-monics. According to voltage harmonic distortion limits set bymarine classification societies, design and sizing of appropriateharmonic attenuation filters is made, including ac and dc chokesand frequency-tuned passive filter options. The effectiveness of theproposed power-quality analyzing procedure is evaluated througha real practical example, which includes harmonic filter designfor VFD retrofit application to fan and pump motors that operateconstantly during sea-going operation in a typical tanker vessel.Power-quality field measurements obtained through a harmonicmonitoring platform implemented on board verify that total volt-age harmonic distortion and individual voltage harmonics at PCCare maintained below 5% and 3%, respectively, showing thatdesign complies with relevant marine harmonic standards even inthe worst operating case.
Index Terms—AC–AC power conversion, harmonic distortion,industrial power system harmonics, industrial power systems, ma-rine vehicle power systems, power quality, variable-speed drives.
I. INTRODUCTION
IN recent years, the strive for reduction of fuel consump-tion and improvement of energy management has led to
significant increase in variable-frequency drive (VFD) use onboard ships and in offshore installations for power conditioning
Manuscript received March 6, 2014; revised June 17, 2014; accepted July 21,2014. Date of publication August 13, 2014; date of current version March 17,2015. Paper 2013-PSEC-0998.R1, presented at the 2014 IEEE/IAS Industrialand Commercial Power Systems Technical Conference, Fort Worth, TX, USA,May 20–23, and approved for publication in the IEEE TRANSACTIONS ON
INDUSTRY APPLICATIONS by the Power Systems Engineering Committeeof the IEEE Industry Applications Society. This work was supported byThenamaris Ships Management Inc.
S. V. Giannoutsos is with Thenamaris Ships Management Inc., 16671Athens, Greece, and also with the Department of Electrical and ComputerEngineering, National Technical University of Athens, 15780 Athens, Greece(e-mail: [email protected]; [email protected]).
S. N. Manias is with the Department of Electrical and Computer Engineer-ing, National Technical University of Athens, 15780 Athens, Greece (e-mail:[email protected]).
Color versions of one or more of the figures in this paper are available onlineat http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TIA.2014.2347453
and energy-saving purposes [1]–[6]. However, the introductionof such nonlinear loads in the vessel’s power system inducesharmonic currents that are often drawn from sources of limitedshort-circuit capability, such as diesel generators (D/Gs), result-ing in high levels of voltage harmonic distortion if left untreated[7]–[11]. It is a known issue that combination of harmonicdistortion and commutation notching can cause hunting andinstability in voltage and frequency regulation control loopsused by D/G electronic governors, let alone the creation ofresonance in power systems where capacitors are installed forpower-factor correction purposes [12], [13].
The importance of using frequency-based harmonic indexesto assess propagation of disturbances in the system and estimatevoltage distortion is demonstrated in several recent studies[14]–[16]. Power-quality problems have become important is-sues to be considered, leading to the recommended limits ofharmonics in supply current standard by IEC, 61000-3-2 [17].Moreover, IEEE Standard 1159 provides the recommendedpractices for defining, measuring, quantifying and interpretingelectromagnetic disturbances in the power system [18]. In[19] and [20], such assessments were performed in shipboardsystems. To address concerns associated with acceptable elec-trical power system distortion on ships and offshore rigs andplatforms, new standards were adopted, and limits were definedto ensure adequate quality and reliability in marine industry[21]–[23].
To this end, studies on power-quality improvement involvingharmonic mitigation, reduction in deviation of voltage fluctua-tion, and unbalance were performed in [24] and [25], whereasplanning studies for harmonic filtering were presented in [26].The influence of reactances to compensate for input currentharmonics generated by VFDs was studied in [27], whereasin [28] and [29], passive harmonic filters were designed toenhance power quality in the shipboard system. The aboveapproaches, however, do not define an effective harmonic an-alyzing procedure in relation to the vessel’s autonomous powersystem in order to verify compliance with harmonic standardsimposed by marine classification societies when VFDs areapplied to motors on board, while ensuring vessel’s powersystem stability.
In this paper, a systematic power-quality assessment andmonitoring methodology is proposed to define the size andtype of harmonic filters that should be used when VFDs are
1910 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 51, NO. 2, MARCH/APRIL 2015
Fig. 1. Effects of harmonic currents on power system impedances. (a) Powersystem impedances between nonlinear load and power source. (b) Voltage dropsassociated with harmonic currents injected to the grid.
applied to motors on board marine vessels in order to com-ply with specified marine harmonic standards. Based on thespecific information related to the nonlinear load nature andshort-circuit capacity of the source, the contribution to voltagedistortion by the VFDs at the point of common coupling (PCC)is calculated. According to existing harmonic distortion levelsin the ship’s power system, appropriate selection and sizing offilters is made. To monitor the effectiveness of the approach,a power-quality monitoring platform is implemented on boarda typical tanker vessel, where harmonic filter design is madefor VFD application to engine room (E/R) ventilation fanand cooling sea-water pump (C.S.W. P/P) motors for energysaving purposes. Experimental results are obtained from on-board power-quality field measurements performed before andafter VFD retrofit installation. It is verified that total voltageharmonic distortion (THDv%) at PCC remains below 5% andthat individual voltage harmonic magnitudes up to the 25th aremaintained below 3% of fundamental frequency magnitude inall ship operating conditions, complying with the AmericanBureau of Shipping (ABS) Rules and Regulations [30].
II. HARMONIC ATTENUATION METHODS
FOR MARINE VESSELS
A. Common Sources of Harmonics for Marine Vessels
Voltage harmonic distortion appears when a nonlinear loaddraws distorted current from the supply, passing through all ofthe impedances between load and power source, as presented inFig. 1(a). The associated harmonic currents cause voltage dropsfor each harmonic frequency across system impedances, asshown in Fig. 1(b). The vector sum of all the individual voltagedrops results in total voltage harmonic distortion (THDv),the level of which depends on system impedance and on thenonlinear load being in service at any specific moment.
In particular, the voltage drop at the load side Vn, load at anygiven frequency can be calculated as
Vn, load = In × (Zc,n +Xt,n +Xd,n) (1)
where In is current at nth harmonic, and Zc,n, Xt,n, and Xd,n
are cable, transformer, and source impedances, respectively.In comparison to shore-based utility power supplies, ef-
fects of harmonic voltages and currents are significantly morepronounced on D/Gs onboard vessels due to their source
Fig. 2. Common harmonic injecting consumers in marine vessels. (a) Switch-mode power supply topology for electronic equipment. (b) VFD topology andharmonic filter options.
impedance being typically three to four times that of utilitytransformers. Therefore, their major effects on D/Gs includelocalized heating and torque pulsations caused by pairs ofpositive–negative harmonic components, thermal and copperlosses, generation of eddy currents, and governor hunting andautomatic voltage regulator (AVR) instability when they occurin combination with ringing or line notching effects.
Common nonlinear loads that inject harmonics in the vessel’selectrical network are switch-mode power supplies and VFDs,the common topologies of which are presented in Fig. 2. Asshown in Fig. 2(a), single-phase full-wave diode rectifiers areused mainly in power supplies of marine consoles, injectingharmonic currents based on two rectified current pulses percycle (p = 2), whose order is defined as
The magnitude of each harmonic component is determined bythe nature of the load, whereas its effect on voltage distortionpresent at PCC depends on the source reactance.
Considering that frequency control is often applied to largemotors, total current harmonic distortion of supply current(THDi) can reach 84% if no harmonic mitigation method isapplied [8]–[10]. Thus, to maintain voltage harmonic distor-tion within permissible limits, attenuation measures have to
GIANNOUTSOS AND MANIAS: ASSESSMENT AND DESIGN METHODOLOGY FOR VFD 1911
be applied. Several VFD harmonic filtering options shown inFig. 2(b) are available depending on the application and desiredlevel of attenuation, including ac-line and dc-link reactorsand ac-side passive filters, which are connected in parallel tononlinear loads and are “tuned” to lower system impedance atthe harmonic frequency desired to be mitigated.
B. Evaluation of VFD Harmonic Mitigation Options for Ships
For the VFD topology shown in Fig. 2(b), which is com-monly used in retrofit installations to motors onboard marinevessels, three types of harmonic attenuation measures will beinvestigated in terms of performance and relative cost to thedrive itself: ac and dc reactances and frequency-tuned passivefilters. While VFD options with a higher number of pulses, oractive front ends, exist, which can push harmonics to muchhigher order, their applications on board vessels are very limiteddue to lack of available cost-effective market solutions withmarine-type approvals from classification societies.
1) Option 1—AC-Line Reactance: AC-line reactors (Lac)can be used primarily as filters to slow down the rising rateof current (di/dt) by limiting the rate of voltage rise whencircuit supply conditions create a voltage step change due toinduced voltage across its terminals. With reference to thecurrent flowing in the ac supply, ii,a, and to the line-to-linesupply voltage, vLL, the required filter inductance L can becalculated according to the desired voltage drop across thereactor vL,ch at fundamental frequency f , as expressed in thefollowing:
L[H] =VL,ch
2πf · Ii,a=
ΔνL,ch% ·(
VLL√3
)100 · 2πf · Ii,a
(4)
where parameter ΔvL,ch%, often referenced by drive and reac-tor manufacturers, is defined as the percentage ratio of desiredvoltage drop across the reactor to the drive input phase voltageat fundamental frequency, as expressed in the following:
ΔνL,ch% = 100 · VL,ch
VLL√3
. (5)
Based on (4), the required filter inductance L can be calculatedfor any value of desired voltage drop ΔvL,ch%. The effects ofvarying ac-line inductance value to current harmonic compo-nents injected back to the grid are investigated for the VFDtopology presented in Fig. 2(b). The variation of input currentharmonic magnitudes Iia,n% for n = 5, 7, . . . , 19, as percent-ages of fundamental frequency magnitude, with ΔvL,ch% forac-line reactance application is presented in Fig. 3.
The results are presented for VFD application to a 75-kW450-V 60-Hz induction motor for a typical 13.4% D/G sourcesubtransient reactance X ′′
d , resulting in Ld = 80 μH. As shownin Fig. 3, harmonic current levels are higher when ΔvL,ch%associated with the ac-line reactor is less than 1% and decreaseswith increasing ac reactance value, with 5% being its maximumrecommended design value due to undesirable high voltagedrop across the reactance.
Fig. 3. Variation of VFD input current harmonic magnitudes Iia,n% for n =5, 7, . . . , 19 with percentage voltage drop across ac-line reactance, ΔvL,ch%.
Fig. 4. Variation of VFD input current harmonic magnitudes Iia,n%, forn = 5, 7, . . . , 19 with equivalent percentage voltage drop across dc choke,ΔvL,ch%.
2) Option 2—DC Reactance (Choke): To avoid voltagedrops associated with ac-line reactors, chokes can be installedin the dc link of VFDs (Ldc), with their values calculated fora desired equivalent voltage drop ΔvL,ch% as in (4). However,these drives normally need discrete surge suppression to protectthe input bridge rectifier devices and limit surges that couldaffect the dc-bus voltage levels. To this end, the rated currentof the dc choke, Idc,av, must be the same as the output currentof the diode rectifier, as shown in the following:
Ii,a =Idc,avcosφ
(6)
where cosϕ is the power factor of the power source, and Idc,avis the average value of dc-bus current, given by
Idc,av =PM(
3√2
π VLL
)· ηM · ηinν
∼= PM
1.35VLL · ηM · ηinν(7)
where PM is motor power, and ηM and ηinv are motor andinverter efficiencies, respectively,
The variation of input current harmonic magnitudes Iia,n%for n = 5, 7, . . . , 19, as percentages of the fundamental fre-quency magnitude, with ΔvL,ch% for dc-choke application inthe VFD dc-link is presented in Fig. 4. It is noted that thepercentage of harmonic currents below rated load will be higher
1912 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 51, NO. 2, MARCH/APRIL 2015
Fig. 5. Variation of the fifth-order current harmonic magnitude, Iia,n%, whena combination of ac-line and dc-bus reactances are used.
than anticipated because dc reactors are usually designed topartially saturate at rated load.
On larger drives, both ac-line and dc-bus reactors may beused to further attenuate harmonics when supply is suscep-tible to disturbances. Simulation results that are presented inFig. 5 show the magnitude variation of the fifth-order currentharmonic magnitude when a combination of ac-line reactancesand dc chokes is used.
3) Option 3—Frequency-Tuned Passive Filters in AC Side:Another option for VFD input current harmonic mitigation is touse passive filters in the ac side, which are connected in parallelwith the nonlinear load and tuned to lower system impedanceat a selected resonant frequency. These filters consist of bothreactors Lf and capacitors Cf , as shown in Fig. 2(b). However,the inductance of the source Ld has to be considered due tothe production of parallel resonance at lower frequency thanin the case of series resonance, possibly leading to powersystem positive feedback and thus to potential misfire of powersemiconductor switching devices.
Based on Fig. 2(b), the total simplified system impedanceZn, at a given harmonic order n, is given by
Zn = jωn × (Ld + Lf ) +1
jωn × Cf
= j
[2πnf1 × (Ld + Lf )−
1
2πnf1 × Cf
](8)
where f1 is the fundamental frequency, and ωn is the fre-quency at harmonic order n in radians per second. The absoluteimpedance can be calculated in ohms over a range of frequen-cies via
|Zn| =∣∣4π2n2f2
1 × (Ld + Lf )− 1∣∣
2πnf1 × Cf. (9)
The components of the passive filter can be selected in order tobe tuned in a specific harmonic order n through
n =fnf1
=
√Xc
XL, tot⇒ (Ld + Lf )× Cf =
1
4π2n2f21
. (10)
Fig. 6. Absolute impedance characteristics for tuned seventh-harmonic pas-sive filter connected in parallel with nonlinear load (Tuned passive filter withLf,7 = 2.44 mH, Cf,757 μF, Qc = 12.5 kVAr).
Fig. 7. Power system absolute impedance variation for multilimbed passivefilter tuned at 5th, 7th, 11th, and 13th harmonic components.
Thus, the defined filter resonates at a frequency as in thefollowing:
fo =1
(2π√
(Lf + Ld)× Cf ). (11)
For example, the attenuation of the seventh harmonic requiresthe filter to resonate at 420 Hz, and according to (10), if acapacitor is selected at 57 μF with reactive capacity of Qc =12.5 kVAr at 440 V and 60 Hz, the equivalent reactance shouldthen be Lf = 2.44 mH, considering that Ld = 80 μH as before.For this case, Fig. 6 presents absolute impedance variation fora seventh-harmonic tuned filter over a range of frequencies.
In Fig. 6, it is shown that the selection of a harmonic tuningfilter at seventh harmonic creates parallel resonance at fifth har-monic and increases the value of system impedance at 300 Hz,resulting in additional voltage distortion caused by VFD re-spective harmonic currents. That parallel resonance can beshifted in a lower frequency, either by connecting resistancesin parallel or by adding another fifth-harmonic tuned passivefilter. To this end, Fig. 7 shows the system impedance variationif a multilimbed harmonic filter tuned in 5th, 7th, 11th, and13th harmonics is used instead. The latter practice can beused on industrial applications and may provide a THDi% of14%–18%. Nevertheless, if such filters are installed in marinevessels, where D/Gs operate in an autonomous power system, itcan lead to potential issues since D/Gs cannot withstand morethan around 20% leading kVAr impressed from passive filter tothe source due to potential armature winding reaction, resultingin over excitation and AVR instability. Moreover, their design
GIANNOUTSOS AND MANIAS: ASSESSMENT AND DESIGN METHODOLOGY FOR VFD 1913
TABLE IPERFORMANCE COMPARISON OF VFDHARMONIC ATTENUATION METHODS
TABLE IIRELATIVE COSTS OF VFD HARMONIC FILTERING SOLUTIONS FOR SHIPS
can be more complicated due to frequency variations presentin the vessel’s power network. Therefore, such filter options invessels are recommended only if in-depth harmonic studies andback-up measurements are performed.
Based on the above results and practices, the effects of VFDharmonic attenuation methods recommended for the marineindustry are summarized and presented in Table I. The currentharmonic percentages In% provided for each filtering solutionwill be used as parameters in the power-quality assessmentmethodology described in the succeeding section, leading toproper filter selection and sizing when VFDs are applied in ma-rine vessels in order to comply with marine harmonic standards.The relative costs of each harmonic filtering solution in relationto VFD unit cost itself Cref,VFD are presented in Table II.
III. PROPOSED POWER QUALITY
ASSESSMENT METHODOLOGY
During the design stage of VFD application to electricmotors on board marine vessels, a harmonic study must beperformed to assess the levels of voltage distortion and toensure that it complies with the requirements of marine classifi-cation societies. In particular, according to Part 4, Ch.8, Sec. 2,par. 7.21 of ABS Rules, THDv% must not exceed 5% at PCCand the magnitude of any individual harmonic component upto 25th must not exceed 3% of fundamental voltage magni-tude [30].
The proposed harmonic analysis procedure and monitoringmethodology are presented in this section with reference to atypical ship electrical distribution topology, as shown in Fig. 8,for the case where VFDs are applied to E/R ventilation fans andC.S.W. pump motors for energy saving purposes. It is shownthat three identical D/Gs with the same kVA rating Si, loadcurrent IL,i, and subtransient reactance X ′′
d,i are connected to
Fig. 8. Power distribution system for a typical ship—Application of proposedharmonic assessment and monitoring methodology for VFD retrofit installation.
a common 450-V bus bar, where the VFD applied motors arealso connected, determining PCC.
The kVA rating of one D/G is chosen as system base refer-ence (Sbase = Si). Since two D/Gs with subtransient reactanceX ′′
d,i usually operate in parallel, the equivalent subtransientreactance should be calculated as in the following:
X∗d_eq[%] =
X ′′d1 ×X ′′
d2
X ′′d1 +X ′′
d1
. (12)
The power-network short-circuit capacity at PCC Ssc is thencalculated for the combined system as follows:
Ssc[kVA] =Sbase
X ′′d_eq
. (13)
The short-circuit current at PCC Isc can then be calculated asfollows:
Isc[A] =Ssc√
3× VLL
= 2× IL1
X ′′d1
. (14)
The equivalent D/G reactance is calculated in ohms, basedon equivalent short-circuit capacity at PCC, as shown in thefollowing:
Xd[Ω] =V 2LL
Ssc. (15)
The total input current of harmonic injecting loads is calculatedfor the motors where VFDs are applied to, as shown in thefollowing:
IL_tot[A] = ILtot_PP + ILtot_fan
=NPP × IL_PP,i +Nfan × IL_fan,i. (16)
The total input load harmonic current values for all harmonicorders up to 25th can be calculated with reference to theharmonic percentages presented in Table I, In%, according
1914 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 51, NO. 2, MARCH/APRIL 2015
to the filtering method used for each drive, as shown in thefollowing:
(18)Magnitudes of existing power system voltage harmonicsVLLn,sys are obtained before VFD application by power-quality measurements; therefore, the total generated voltageharmonics are
The equivalent percentages of harmonics with reference to line-to-line voltage are tested for compliance to specified marineclassification society limits, as in the following:
%VLLn,tot = 100%× VLLn,tot
VLL
, n = 5, 7, 11, . . . , 25.
(20)
The total voltage harmonic distortion (THDv%) can then becalculated at PCC and checked against related requirements asfollows:
THDv%
= 100%×
√25∑
n=5V 2LLn,tot
VLL
=
√V 2LL5, tot + V 2
LL7, tot + V 2LL11, tot + · · ·+ V 2
LL25, tot
VLL
.
(21)
Since calculations are made for worst-case scenario, whereVFD equipped motors run at 100% of nominal speed (60 Hz),the values of requested harmonic indexes should be belowlimits during the whole vessel’s operational profile. If systemdoes not comply with the required limits, a different harmonicmitigation method from the ones described in Table I should beconsidered, and calculations have to be repeated.
To ensure that power quality at PCC is always within spec-ified limits, a harmonic monitoring platform with real-timedata acquisition features is implemented on board, as shownin Fig. 8. It consists of analyzing modules that obtain real-time voltage and current measurements after the molded casecircuit breakers (MCCB), connecting the respective VFD tomain switchboard. The modules communicate with a wirelessaccess point using Ethernet over TCP/IP. The data acquisitioncomputer acts a time server, synchronizing in real-time withthe modules. This allows logged data from multiple units to bedisplayed on the same time scale, enabling root-cause analysisby processing conditions leading up to, during and after events.
Fig. 9. Onboard power-quality measurements before VFD application.(a) Line-to-Line voltage waveform at PCC. (b) Harmonic spectrum showingindividual harmonic magnitudes and THDv% of measured voltage waveformat PCC.
IV. CASE STUDY—VFD INSTALLATION
IN A TANKER VESSEL
The proposed methodology is applied during the design stageof VFD retrofit installation to all four 18.5-kW 33.6-A 450-V60-Hz E/R ventilation fans and to one 75-kW 129.1-A 450-V60-Hz C.S.W. P/P on board an Aframax size tanker vessel withreference to the topology presented in Fig. 8. The efficiencyfactors for fan and pump motors are 0.9 and 0.95, respectively.DC reactances of 3% are used for VFDs, which are applied toE/R fans, and a 3% ac–dc reactance combination is selected forVFD application to C.S.W. pump. According to (4) and (5), the3% reactance used in the dc link of E/R fan drives is definedat 600 μH, whereas the 3% ac and dc reactances used inthe C.S.W. pump drive are 150 μH each. These motors areconstantly used during vessel’s operation and were previouslystarted D.O.L; therefore, a harmonic study is required duringthe design stage to verify compliance with specified limits atany moment.
The power-quality measurements performed at PCC beforeVFD application on board are presented in Fig. 9. From theprovided harmonic spectrum of voltage waveform, it is shownthat existing power system voltage harmonics include fifth- andseventh-harmonic components with magnitudes VLL5, sys =
3 V and VLL7, sys = 2 V, respectively, whereas THDv% beforeVFD application was 0.7%, mainly due to marine consolepower supplies, which use one-phase or three-phase dioderectifiers.
Table III shows data and calculated parameters related to thevessel’s power system with reference to (12)–(16) and Fig. 8.The total magnitudes of input current harmonics generated byVFD operation at nominal speed (60 Hz) are calculated forharmonic orders up to 25th and are presented in Table IV.
The corresponding harmonic voltages created due to voltagedrop on the power system impedances are also presented inTable IV with reference to (17) and (18) and to the selectedharmonic filtering options as described in Table I.
The total magnitudes of the voltage harmonics up to the25th order present at PCC, including existing power systemvoltage harmonics as shown in Fig. 9, are presented in Table Vwith reference to (19). The percentages of individual voltageharmonics with respect to the fundamental frequency are thencalculated and are shown to be below the required limits of3%. The total harmonic distortion (THDv%) is also below therequired limit of 5%, verifying proper selection of harmonic
GIANNOUTSOS AND MANIAS: ASSESSMENT AND DESIGN METHODOLOGY FOR VFD 1915
TABLE IIIPOWER SYSTEM PARAMETERS FOR THE MARINE VESSEL UNDER STUDY
TABLE IVVFD INDUCED HARMONICS FOR THE VESSEL UNDER INVESTIGATION
filters for the application. Power-quality field measurements areobtained from the vessel under study after the VFD topologyretrofit installation and are presented in the succeeding sectionto validate the proposed power-quality assessment and filterdesign method.
V. EXPERIMENTAL VALIDATION OF
PROPOSED METHODOLOGY
Following the results from the proposed power-quality as-sessment procedure, the designed harmonic filters are includedin the VFD topology installed as a retrofit to one 75-kW C.S.W.P/P and to four 18.5 kW E/R fan motors in the tanker vesselunder study, as shown in Fig. 10(a) and in Fig. 10(b), respec-tively. Since marine-type armored power cables with copperor steel wire braid are used from the inverters to the motors,while distance is kept below 80 meters, there is no need to
TABLE VSHIP’S POWER SYSTEM VOLTAGE HARMONIC DISTORTION
install common-mode chokes for prevention of inverter-inducedbearing currents [33], [34]. The designed harmonics monitoringplatform is implemented on board to effectively monitor powerquality of voltage and input current waveforms during vessel’ssea-going period. Power-quality measurements are performedat PCC in order to confirm installation compliance to specifiedmarine harmonic standards and verify the effectiveness of thedesign.
As shown in Fig. 10(c), real-time voltage and current mea-surements are performed by ELSPEC G4430 power-quality an-alyzing modules used to monitor line-to-line voltage and inputsupply current waveforms. For current waveform recording,current transformers are installed after the respective MCCBat the starter panel of each motor, as shown in Fig. 10(d).Communication with a router acting as a wireless access pointis performed using Ethernet over TCP/IP. Common FTP pro-tocol is used for data acquisition in a portable computer. Dataanalysis is made using PQSCADA configuration software. Incomparison to other measuring configurations, this topology isable to record and store voltage waveforms at 1024 samplesper cycle and current waveforms at 256 samples per cyclefor more than a year without gaps in the data, making itideal for the marine environment. Since the data acquisitioncomputer is used for time synchronization of the analyzingmodules, it is possible to display and analyze outputs fromdifferent modules at the same time scale. This provides theopportunity to investigate the origin of events that occur fromdifferent sources. The parameter measurements and evalua-tion are performed according to relevant IEC 61000-2-4 andIEC 61000-4-30 standards for industrial networks [31], [32].
1916 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 51, NO. 2, MARCH/APRIL 2015
Fig. 10. Implementation of designed VFD topology and power-quality monitoring platform in the tanker vessel under study. (a) VFD unit with associated 3%dc and 3% ac reactances applied to C.S.W. pump motor. (b) VFD unit with associated 3% dc choke applied to E/R fan motor. (c) Layout and interconnections ofpower-quality analyzer module. (d) Measurement points after respective motor power supply MCCB.
Fig. 11. Power quality analysis of input supply current for VFD pumps and fan motors operating at 40 Hz. (a) PCC Voltage and input supply current waveformfor C.S.W. P/P motor. (b) Harmonic spectrum and distortion of C.S.W. P/P motor input supply current. (c) Input supply current waveform for E/R fan motor.(d) Harmonic spectrum and distortion of E/R fan motor input supply current.
Fig. 12. Onboard field measurements and analysis for variable-frequency-driven C.S.W. P/P motor during typical sea-going operation. (a) C.S.W. P/P motoractive power consumption waveform. (b) C.S.W. P/P motor VFD input current RMS waveform. (c) Variation of individual input current harmonic magnitudes(5th, 7th, and 11th order) as percentages of fundamental frequency magnitude. (d) Total harmonic distortion of input current waveform.
The voltage waveform at PCC and the corresponding inputsupply current waveform for a typical C.S.W. P/P motor op-eration at 40 Hz are presented in Fig. 11(a). To determine thepower quality of the current waveform, its harmonic spectrumis presented in Fig. 11(b). It is shown that fifth- and seventh-order harmonics are prevalent with magnitudes of 21% and8% of fundamental current magnitude, respectively, whereas
THDi% is 23.2%, validating the harmonic filter performancefor this particular application since a combined 3% dc and3% ac choke selection was made. The input supply currentwaveform for typical E/R fan motor operation at 40 Hz ispresented in Fig. 11(c), whereas its corresponding harmonicspectrum is presented in Fig. 11(d). It is shown that prevalentfifth- and seventh-order harmonic magnitudes are 36% and
GIANNOUTSOS AND MANIAS: ASSESSMENT AND DESIGN METHODOLOGY FOR VFD 1917
Fig. 13. Onboard field measurements and analysis of voltage waveform at PCC after the VFD retrofit installation, using the implemented monitoring platform.(a) C.S.W. P/P motor active power consumption waveform. (b) E/R ventilation fan motor active power consumption waveform (same for fans 1–4). (c) Variationof individual voltage harmonic magnitudes (5th, 7th, and 11th order) as percentages of fundamental frequency magnitude. (d) Total voltage harmonic distortionTHDv%.
15% of fundamental current magnitude, respectively, whereasTHDi% is 43.3%. Increased THDi% is justified since only a 3%dc choke was selected for this application. Higher total currentharmonic distortion is expected at partial load due to dc chokebeing designed to partially saturate at rated load.
To experimentally verify that harmonic distortion levels inthe voltage waveform at PCC vary within acceptable limitsduring VFD operation as well as evaluate system performanceduring transients, power-quality field measurements are per-formed during typical vessel sea-going operation, using theharmonic monitoring platform installed on board. Onboard fieldlogging measurements for variable-frequency-driven C.S.W.P/P motor operation are presented in Fig. 12. It is shown thatthe load of C.S.W. pump motor varies from nominal 75 downto 24 kW, depending on the speed reference provided by VFDcontrol system. During this period, the magnitudes of the fifth-order input supply current harmonic varies from 22% to 30%,whereas magnitude of seventh-order input current harmonicvaries from 8% to 5%, with respect to the fundamental fre-quency. The 11th-order harmonics is found to be very low, closeto 2%. As a result, total current harmonic distortion, THDi%,varies from 23% to 30%, as expected according to previous casestudy results. An instantaneous increase in THDi% is also notedduring sudden load transients.
Similarly, onboard field logging measurements and analysisof voltage waveform at PCC after the VFD retrofit installationare presented in Fig. 13. During operation, all VFD equippedmotors usually operate at 40–50 Hz; therefore, in Fig. 13, it isshown that the load of sea water cooling pump varies from 50to 24 kW, whereas the load of E/R fans varies from 7 to 2.5 kW.During logging period, it is shown that fifth voltage harmonicat PCC has an average percentage value of 1.5% of nominalvoltage at fundamental frequency or 6.7 V in absolute values.Significantly lower values are recorded for 7th- and 11th-ordervoltage harmonics. If these values are compared with those in
Fig. 14. Voltage power-quality analysis at PCC for VFD operation at 60 Hz(five VFD equipped motors in service). (a) Line-to-Line voltage waveform.(b) Harmonic spectrum and THDv% of voltage waveform.
Fig. 9 before VFD application, it is shown that the magnitudeof fifth voltage harmonic is increased by 1%, but in any case,it is well below the limit of 3% imposed by classificationsocieties. Regarding THDv%, it is shown that its average valueis maintained at 1.8% at current ship operating profile, whichresults in a 1.1% increase if compared to THDv% recordedbefore VFD application. The spike noticed in the diagramis caused by ballast pump (180 kW) autotransformer startingsequence. Nevertheless, during vessel’s operating period, it isshown that THDv% is also defined well below 5% as requiredby ABS marine classification society [30].
When the vessel sails in areas with hot climates and highS.W temperature or when vessel speed is increased, highercombustion air and cooling water flow are required by M/E andauxiliary machinery to compensate for increased heat dissipa-tion requirements, resulting in operation of all VFD equippedE/R Fan and C.S.W. P/P motors close to maximum 60 Hz. Thespecified THDv% and individual harmonic magnitude limitsmust be met also in this worst operating case. As shown inFig. 14, THDv% is equal to 3.4%, and magnitudes of fifth andseventh harmonics are equal to 2.5% and 1.3%, respectively,
1918 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 51, NO. 2, MARCH/APRIL 2015
with reference to fundamental frequency, verifying the effec-tiveness of proposed design also at this case.
VI. CONCLUSION
In this paper, a power-quality assessment and monitoringmethodology has been introduced to dimension harmonic filtersfor VFD application to electric motors in marine vessels inorder to comply with marine harmonic standards. A highly sys-tematic approach was proposed, considering the importance ofpower system impedances to avoid resonance and D/G governorand AVR instability. The presented methodology is highly pa-rameterized to be adapted for any marine power system facingstructural changes due to VFD application. For the vessel underinvestigation, the proposed power-quality analysis methodol-ogy was used to define the filter design requirements for VFDapplication to E/R fan and C.S.W. pump motors, which areconstantly used during sea-going and cargo handling periods. Apower-quality monitoring platform was implemented on boardto record and analyze current and voltage distortion at PCC.Power-quality measurements are obtained before and after theimplementation of the designed VFD topology in the vesselunder investigation. Experimental results show that the mag-nitudes of individual voltage harmonics up to 25th are below3% of the magnitude of the fundamental frequency, whereasTHDv% is well below 5% even in the case where drives operateall subject motors at 60 Hz, validating the effectiveness of theproposed analysis and design.
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[19] P. Crapse, J. Wang, J. Abrams, and Y.-J. Shin, “Power quality as-sessment and management in an electric ship power system,” in Proc.IEEE Elect. Ship Technol. Symp., Arlington, VA, USA, May 2007,pp. 328–334.
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[21] P. F. Ribeiro, M. Steuer, and M. Islam, “Reevaluating electric powersystem harmonic distortion limits for shipboard systems,” in Proc.Harmon. Qual. Power Int. Conf., Lake Placid, NY, USA, Sep. 2004,pp. 706–711.
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[24] S. X. Duarte and N. Kagan, “A power-quality index to assess the im-pact of voltage harmonic distortions and unbalance to three-phase induc-tion motors,” IEEE Trans. Power Del., vol. 25, no. 3, pp. 1846–1854,Jul. 2010.
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GIANNOUTSOS AND MANIAS: ASSESSMENT AND DESIGN METHODOLOGY FOR VFD 1919
Spyridon V. Giannoutsos (S’12) received theDiploma in electrical and computer engineeringfrom the National Technical University of Athens(NTUA), Athens, Greece, in 2010. He is currentlyworking toward the Ph.D. degree in marine electricalengineering, power electronics, and process controlwith the Electrical Machines and Power ElectronicsLaboratory, NTUA.
Since 2011, he has also been a Research Associatewith the Electrical Machines and Power ElectronicsLaboratory, NTUA. In 2012, he joined Thenamaris
Ships Management Inc., Athens, where he is holding the position of FleetSupport Electrical Engineer, focusing on design and implementation of energyand fuel saving solutions on board. His research interests include processcontrol, power electronics, motor drives, and power quality.
Mr. Giannoutsos is a member of the IEEE Industry Applications and IEEEIndustrial Electronics Societies and a Registered Professional Engineer of theTechnical Chamber of Greece.
Stefanos N. Manias (M’85–SM’92–F’05) receivedthe B.Eng., M.Eng., and Ph.D. degrees fromConcordia University, Montreal, QC, Canada, in1975, 1980, and 1984, respectively, all in electricalengineering.
In 1975, he joined the Canadian BroadcastingCorporation, where he designed radio and televisionautomation systems. From 1979 to 1981, he waswith Northern Telecom Canada where he designedswitching-mode power supplies. Since 1989, he hasbeen with the Department of Electrical and Com-
puter Engineering, National Technical University of Athens, Athens, Greece,where he is currently a Full Professor, conducting research in the areas of powerelectronics and motor drives with more than 80 IEEE and IEE publications.
Dr. Manias is the Chapter Chair and the Founder of the Greece Section of theIEEE Industry Applications (IAS), IEEE Power Electronics (PELS), and IEEEIndustrial Electronics (IES) Societies, and a member of IEEE Motor DrivesCommittee. He is a Registered Professional Engineer in Canada and Europe.
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A data-driven process controller for energy-efficient, variable-speed fan operation in marine vessels
By spyr Idon v. g IAnnoutsos & stEFAnos n. MAnIAs
he capacity of engine room (e/r) ventilation fans installed in
marine vessels is defined during the design stage based on the maximum air flow
required by the main engine (m/e), diesel generators (d/gs), and boilers. however,
the vessel’s operating profile at reduced m/e power and speed does not justify the use
of full capacities. in this article, a data-driven process controller is proposed to adjust the speed of
e/r ventilation fan motors according to the variation of combustion air-flow and heat emission
requirements to optimize energy efficiency during a ship’s seagoing and cargo operation periods. the
dynamics of the adaptive controller are established through parameter monitoring, regardless of the
model of the controlled plant. for the tanker vessel under investigation, a case study initially defines
IMPROVING ENGINE ROOM VENTILATION SYSTEMS
Digital Object Identifier 10.1109/MIAS.2015.2459088
Date of publication: 7 September 2016
T
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the plant operational constraints and its power-saving potential. the proposed variable-frequency process con-trol topology is then applied as a retrofit installation to the existing 18.5-kW e/r fan motor starters. onboard experimental results show significant improvement in the vessel’s power balance, d/g fuel consumption, and level of emission factors [carbon dioxide (co2), sulfur dioxide (so2), and nitrogen oxides (nox)], validating the performance of the proposed control topology, which is implemented using commercial programmable logic controllers (plcs) and variable-frequency drive (vfd) units.
Methods of Fuel Consumption Efficiency With rising fuel prices, tighter regulations, and increasing environmental concerns, fuel consumption efficiency is a high priority for charter boat operators as well as for ship operators in the maritime industry. to optimize energy management, attempts to reconfigure the control of the onboard power system have been made in [1] and [2], while a waste heat recovery system was proposed in [3]. in addition, ships today adopt the slow steaming practice and operate the m/e at lower speeds, leading to varying flow requirements for the machinery associated with water cooling and air ventilation systems due to the need for derated capacities [4].
in [5]–[7], techno-economical studies suggest that the application of vfds to electric motors for flow regulation purposes can lead to significant energy sav-ings. options for vfd applications in marine vessels are provided in [8] and [9], while an energy-saving evaluation method for adjustable-speed-driven sea water cooling pumps is described in [10]. optimal vector control design methods for ventilation and pumping induction-motor systems are modeled and presented in [11] and [12], while a converter-based starting method associated with speed control for cen-trifugal loads has been reported in [13]. detailed analysis and performance evaluation of frequency con-trolled centrifugal fans and pumps subject to voltage variation and unbalance is also performed in [14]–[16]. however, the mentioned model-based approaches are limited to the control of the frequency converter and motor itself, based on sufficient quantitative knowledge being available. this is not often the case, especially in marine vessels where the flow regulation relies on varying process requirements, which depend on a ship’s operating profile.
to ensure reliability and safety in the modern indus-trial processes, data-driven methods have received con-siderable attention, particularly for process monitoring purposes [17]–[19]. these approaches can be catego-rized into pure data based, which include real-time sampling of state signals [20], iterative learning for con-troller tuning [21], and model-free adaptive approaches [22], as well as model-data integrated methodologies, which focus on automatic controller feedback tuning [23]. to this end, a self-tuned, data-driven process con-troller for energy-efficient, variable-speed pump opera-tion in a central cooling water system of marine vessels was designed and implemented in [24]. the results
showed significant improvement in energy and fuel consumption compared to existing approaches.
as an extension of the previous work, a data-driven process control topology is introduced in this article to adjust the speed of fan motors in e/r air ventilation sys-tems of marine vessels according to varying combustion and heat evacuation air-flow requirements to optimize process energy management. Based on proportional-inte-gral-differential (pid) control, the controller parameter adjustment is based on real-time pressure and temperature measurements from the process, which are used as feed-back signals to adapt to process constraints regardless of the controlled plant model. case study results are initially obtained from modeling the vessel’s e/r air ventilation system according to derated m/e and auxiliary machinery combustion and heat emission air-flow capacities, which
are related to slow steaming practices to examine the plant power-saving potential and determine the critical process parameters.
the proposed data-driven control topology for ener-gy-saving optimization of a vessel’s e/r ventilation sys-tem is implemented using a commercial plc, while a human–machine interface (hmi) is included for plant parameter monitoring. the proposed topology is applied as a retrofit installation to the existing e/r fan starters in a typical tanker vessel. experimental results show that controller dynamics are well established, lead-ing to significant energy savings as well as to d/g fuel consumption and emission factors reduction. power quality measurements are also performed to show that voltage distortion at point of common coupling (pcc) remains within acceptable levels required by marine classification societies.
Flow Regulation for Variable-Speed-Driven Fansfor the majority of e/r air ventilation systems installed in ships, the impeller of axial flow fans is supplied with the blades adjusted to the pitch angle b , corresponding to the design operating point. the total pressure increase in pascal ( )Pa pft, developed by an axial fan operating at its design point depends on the air-flow density in kilo-grams per cubic meter t, the rotational component imparted by the impeller uC , and the peripheral velocity of the impeller in meters per second .u however, both velocity components vary proportionally with the rota-tional speed in revolutions per minute n and the impel-ler diameter in meters d as shown in
( ) ( )u C nd nd n d# # # # # #? ?t t t= .p u2 2
ft (1)
the radial air volume flow in cubic meters per second q delivered by the fan at the impeller outlet for an axial impeller is given by
( / ) ( / ) ( )q d C d nd n d4 42 2 3# # #? ?r r= o (2)
since vC , which is the axial velocity in meters per second, is proportional to rotational speed and impeller diameter.
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depending on (1) and (2), the output power in watts, pout, for a fan with efficiency h is proportional to pft and q, as shown in
( ) / ( )
( ) .
P p q n d
n d n d
,f2 2
3 3 5
out ft # # #
# # # #
?
?
h t
t
=
(3)
ventilation fans in ships are dimensioned to provide the required air flow for m/e operation at the maximum capacity ratio (mcr), so the operating point of the system is originally a ,), pA ftA(q as presented in figure 1. this results in continuous nominal power consumption, ,P , ,f Aout as the fan motor is working at nominal speed An . due to
vessel slow steaming practice, however, the m/e often operates at lower speeds, resulting in fewer flow-rate requirements from the process. in that case, variable-fre-quency control can be used to adjust the speed of the fan motor to a lower value, Bn , since moving the operating point from a to B )ft, p BB(q results in a significant power-saving margin as highlighted in figure 1. this occurs because the fan power consumption is reduced with flow
according to affinity laws, as shown in
.PP
nn
qq
, ,
, ,
f A
f B
A
B
A
B3 3
out
out= =` cj m (4)
however, since the fan operation is associated with a spe-cific process whose air-flow requirements vary according to the vessel’s operating profile and m/e speed, an adjustment of Bn simply based on the process modeling is not adequate without putting the process safety and reliability at risk. therefore, the objective of the pro-posed control topology is to operate the e/r ventilation fans at an optimal speed based on the process constraints so that maximum power savings are achieved. the mini-mum motor speed limit min~ must be defined according to the minimum required air volume flow rate qmin at each period.
the combustion and heat emission air-flow require-ments of the m/e and auxiliary machinery at reduced speeds can be defined through modeling the vessel’s e/r air ventilation system, which will reveal the maximum power-saving margin that can be obtained after the appli-cation of the proposed topology. the following results are provided specifically for the tanker vessel under study, assisting the design and implementation of the proposed process control scheme.
E/R Air Balance Calculation for the Ship under Studythe typical e/r ventilation system for a marine vessel is presented in figure 2. four axial flow fans are installed outside the engine casing and, through separate air ducts, provide combustion air to major e/r consumers, which include the m/e, d/gs, and the oil-fired boiler. two of them can also work in exhaust mode, used only on very hot days when the air needs to be evacuated from the e/r during hot work. it is noted that air ventilation ducts are placed near m/e and d/g turbochargers.
as per iso 8861:1998 standard [25], the designed flow-rate capacity for the fans q ,tot MCR depends on the required combustion air flow, qc, and the heat exhaustion flow, qh, defined as per
( , ) ( , . ).max maxq q q q q q1 5, , , c h ctot MCR tot I tot II #= = + (5)
the combustion air flow in cubic meters per second, ,qc is calculated for m/e and d/g oper-ation at point N equal to the mcr as in
,q q q q, ,c N N b/ /M E D G= + + (6)
where q ,NM/E is the combustion air flow for the m/e in cubic meters per second, N = 100%; ,ND/Gq is the
combustion air flow for d/g in cubic meters per second, N = 76%; and qb is the air flow for combustion for boilers in cubic meters per second. the required fan air flow for evacua-tion of heat emission in cubic meters
Fan
Tota
l Pre
ssur
e –
p ft(
Pa) pft, A
pft, B
Power Saved by SpeedRegulation (A to B)
Flow Reduced by
Speed Adjustment
(A to B)
B
A
η1%
η1%
η2%
η2%
β4β3
β2
β1
β4
β3β2
β1
nA
O qmin qB qA
Air Volume Flow Rate – q (m3/s)
The performance curves for variable-speed axial flow fans.1
M/E
D/G
Oil-FiredBoiler
Air
Duct MotorFan
Outside
Area
E/RArea
E/R Ventilation System(Number One of Four Ventilation Ducts)
qin
qh
qC
qb
qD/G
qM/E
The layout of a typical E/R ventilation system in marine vessels.2
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per second hq is based on maintaining the increase of e/r air temperature T3 at 12.5 K as expressed in
. ( ),qc T
q q q0 4 , ,hi
N N b/ /M E D G# #
#t D
U= - + +/ (7)
where the total heat to be evacuated includes heat emis-sions (in kilowatts) from the m/e, d/gs, boiler, steam and condensate piping, alternators, electrical equipment, and others as in
., ,i N N b p el/ /M E D G alt othU U U U U U U U= + + + + + +/
(8)
the required combustion air flows for the m/e, d/g, and boiler as a function of their operating capacities can be calculated for NM/E= 100% and ND/G=76% in
( ) /q P m, ,N N ad/ /M E M E # t= (9) ( ) /q P m2, ,N N ad/ /D G sets D G# # t= (10) ( ) / ( , ),q Q m m f 3 600b B fs af s# # # # t= (11)
where P ,NM/E is the m/e load at the operating point N in kilowatts, P ,ND/G is the d/g load at the operating point N in kilowatts, mad is the air requirement for diesel engine combustion (0.0023 kg/kWs for two-stroke engines and 0.0020 kg/kWs for four-stroke engines), t is the density of air (1.13 kg/m3, in 35 °c and 101.3 kpa), QB is the total steam capacity for boilers in kilograms per hour, mfs is the fuel consumption in kilo-grams of fuel per kilogram of steam (typically 0.077 kg/kg), maf is the air requirement in kilograms of air per kilograms of fuel (typically 15.7 kg/kg), and fs is the load factor on tank cleaning condition (typically 0.911 based on steam consumption).
similarly, heat emissions can be expressed as in
. P0 141, ,.
N N0 76
/ /M E M E#U = (12). ( )P f2 0 396, ,
.N N l
0 7/ /D G sets D G# # #U = (13) ( / , ) ( / )Q m h h B f3 600 100b B fs B s1# # # # #U D=
(14) ( / )m h f100p sc p s# #U D= (15) sets ( / )P f2 1 100,alt N l/D G# # #hU = - (16) % ,P f15 ,el N l/D G# #U = (17)
where h is the lower calorific value of fuel (40,200 kJ/kg), Bh3 is the boiler heat loss in percentage at the mcr
(0.37%), B1 is the constant related to boiler location in the e/r (0.1), msc is the steam consumption in kilowatts (1 kW = 1.6 kg/h steam), hq3 is the steam heat loss as
consumption percentage (0.2%), h is the alternator effi-ciency in percentage, fl is the load factor on the tank cleaning condition (typically 0.58 based on the load bal-ance study), and othU is the other heat emissions (e.g., from hot tanks) in kilowatts.
according to (5), the resulting capacity of the ventila-tion system, ,q ,tot MCR should be the maximum value
between the sum of the air required for combustion and heat evacuation and the combustion air corresponding to 1.5 times the total air consumption for the m/e, d/g, and boilers operating at the mcr. the designed air-flow capacity for each fan, ,q ,MCRfan is then defined based on the number of fans, ,Nfan as
/ .q q N, ,fan MCR tot MCR fan= (18)
for the vessel under investigation, table 1 shows the e/r ventilation system-designed capacities for an m/e operating at 100% of the mcr, two d/gs operating at 76% of the mcr, and a boiler operating at nominal capacity. it is shown that four e/r ventilation fans are
selected with a 50,000-m3/h flow capacity coupled with 18.5-kW motor each to compensate for the maximum total required air flow of 200,000 m3/h as calculated by (5), considering combustion and heat emi s -sion needs.
due to slow steaming practices during a seagoing period, however, the m/e and auxiliary machinery are operating at a reduced power and speed far from the mcr point, resulting in lower combustion and heat emission air-flow requirements, which vary with the ship’s operating profile. more specifically, a vessel’s m/e is typically working at derated point N equal to about 60% of the specified mcr due to slow steaming practic-es [4], i.e., at an m/e load of 8,568 kW or at 88.6 r/min propeller speed. Based on (5)–(8), the reduced com-bustion and heat emission air-flow requirements are cal-culated and presented in table 2 for seagoing, port/anchorage, and unloading conditions, based on a vessel’s operating profile.
from this case study, it is shown that due to slow steaming, the air-flow requirement is now reduced to 100,700 m3/h, but with the existing topology, each fan motor still operates at a nominal power of 18.5 kW. if variable-frequency control is used to adjust the speed of each fan, theoretically its produced flow could propor-tionally be reduced to 50% of ,q ,MCRfan resulting in fan
power consumption, pout, of just 2.3 kW, according to (4). however, this practice alone does not satisfy the crit-ical process parameters and constraints that vary among seagoing, anchorage, loading, and unloading operating conditions. in particular, overpressure must be achieved in the e/r for m/e and d/g turbochargers to operate satisfactory; a specific level of the e/r temperature must
be maintained, while minimum speed limits must be introduced, considering the number of fans working at each time. the proposed data-driven control topology, presented in the next section, adapts to the vessel’s operat-ing profile to minimize fan power consumption as well as satisfy dynamically changing process parameters without relying on the process model.
Proposed Data-Driven Process Control Topologythe proposed data-driven controller interface with the e/r air ventilation system process for the vessel under study is presented in figure 3, using real-time
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measurement sets to regulate parameter set points and set fan motor speed.
Controller Design and Operating PrinciplesWith reference to figure 3, pressure in the e/r area PE/R as well as ambient pressure of the area outside the engine cas-ing pout are measured through a differential pressure sensor and provided as control input to the proposed process con-troller. this variable is used to maintain a desired differen-tial pressure reference ,P ,diff E/R
) which has a direct impact on the m/e and d/g turbocharger operation, especially during seagoing period. in addition, a certain level of the e/r ambient temperature T ,amb E/R
) is maintained by providing a relevant input to the controller T ,amb E/R to ensure a suitable temperature level in the m/e scavenge air chamber and to limit the condensation effect. a minimum speed limit min~ is introduced based on the number of operating fans Nfan and on the process physical constraints.
the proposed data-driven process controller layout and interface with the variable-frequency control topology are presented in figure 4 with reference to the vessel studied. Under any circumstances, the desired P ,diff E/R
) and T ,amb E/R
)
set points are achieved through two independent pid con-trol loops, which are used to regulate motor speed references ,P diff~
) and , ,T E/R~) respectively, to satisfy both
the combustion air-flow and heat evacuation air-flow requirements:
( )K e t#+
( ) ( )
K e t K e t dt
dtd
, , , , ,
, ,
*P P P I P
t
D P
0
diff diff diff diff diff
diff diff
# #~ = + #
(19)
( ),K e t#+
( ) ( )
K e t K e t dt
dtd
, , , , ,
, ,
*T P T P T I T P T
t
D T P T
0
/E R # #~ = + #
(20)
where ( )te ,P diff and ( )e t,P T are expressed by
( ) ( )e t P P t, ,*
,P / /diff diff E R diff E R= - (21)
( ) ( ) .e t T T t, ,*
,P T / /amb E R amb E R= - (22)
to comply with dynamic system changes and process con-straints, the output speed command signal from the pid control loops ,T DP~
) is the maximum value between the two references, while the final speed command provided to each fan auto~
) is lower bounded by min~ , as shown in
, , %.max 95,*
,*
,* *
minT P T /DP diff E R auto# #~ ~ ~ ~ ~= " ,
(23)
the pid control tuning occurs independently for each loop by obtaining the ultimate gain ,Ku which the output oscillates at constant amplitude with period
.Tu the calculation of the controller parameters is performed through the expressions described in
. , , .K K KTK K K T0 60 2
8P u Iu
PD
P u# # #= = = (24)
Regulation of Controller Parameters and Set Pointsfrom the case study in the section “e/r air Balance cal-culation for the ship Under study,” it was shown that
TABLE 1. THE E/R VENTILATION SYSTEM-DESIGNED AIR-FLOW CAPACITIES FOR THE M/E MCR OPERATION.
Parameter Description Symbol value Comments
M/E model M/E 6s60Mc-c8.2-tII Maker MAn B&w
Engine load at the MCR point PMCR [kw] 14,280 At MCR, M/E works at
Engine speed at the MCR point UMCR [r/min] 105
D/G capacity PD/G [kw] 780 %Nm/e=100%, and D/Gs work
Oil-fired boiler designed capacity QB [kg/h] 25,000 al %ND/G=76%
required air flow for combustion qc [m3/h] 133,340 Refer to (6)
required air flow for heat evacuation qh [m3/h] 54,220 Refer to (7)
Air flow as per ISO 8861:1998—Case 1 qtot,I [m3/h] 187,560
Air flow as per ISO 8861:1998—Case 2 qtot,II [m3/h] 200,000 Refer to (5)
Total designed air flow (ISO 8861:1998) q ,MCRtot [m3/h] 200,000
number of E/r ventilation fans Nfan 4
E/R ventilation fan designed air flow ,q ,fan MCR 50,000 Refer to (18)
E/R ventilation fan motor power Pfan [kw] 18.5
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air-flow requirements by e/r consumers vary between vessel seagoing, anchorage/loading, and unloading modes, requiring controller set-point adaptation to e/r needs.
Supply Air Pressure Regulationin all operating conditions, the objective is to maintain e/r overpressure throughout e/r space, as shown in
.P p P P 0, / /diff E R f E R out 2= = -t (25)
this overpressure should be slightly positive but not higher than 5-mm water column or 0.5 mbar above the outside pressure at the air outlets in the funnel. the m/e auxiliary blowers operate only during maneuver-ing; during a seagoing period, the m/e turbocharger alone is in service. thus, the level of e/r overpressure set point during seagoing is selected so that the m/e turbocharger can take in pressurized air, and scavenge air pressure P ,M/E scavenge inside the m/e scavenge air chamber is maintained always higher than 0.7 bar to keep the m/e running efficiently. as shown from table 2, during anchorage or cargo operation, the m/e is not working, so pressurized air is needed only by the d/g turbochargers and oil-fired boiler, resulting in fewer combustion needs. therefore, the level of P ,diff E/R
) set point and the minimum allowed motor speed limit
min~ are dynamically selected in the range of 15–45 pa and 50–85%, respectively, according to the number of fans Nfan operating at each period to minimize fan power consumption within the acceptable process limitations.
since the outside air barometric pressure is not steady but varies in the range 980–1,040 mbar, the differential pressure sensor uses a diaphragm, which vibrates according
to difference in pressure measured through tubes between the e/r and outside to provide accurate inputs to the con-troller. the sensor itself is installed in the e/r casing close to the funnel door, where the air-flow direction is stabilized.
E/R Ambient Temperature Regulationmeasurements show that the e/r ambient temperature T ,amb E/R is normally 10–12 °c higher than the ambient outside air temperature T ,amb out, which is relevant to sea
water temperature T .S W [26]. Based on the above, the average temperature in a ventilated e/r is not higher
than the one
TABLE 2. THE AIR BALANCE CALCULATION AT SLOW STEAMING OPERATING POINT N M/E WORKING AT 60% OF THE MCR AND D/GS WORKING AT 76%.
Parameter Description Symbol value Comments
Derated engine load at point N PM/E,N [kw] 8,568 Reference to M/E Project
guide for the specific vesselDerated engine speed at point N nN [r/min] 88.6
Percentage engine load at point N %PM/E,N [kw] 60
Percentage engine speed at point N nN [%] 84.3
required air flow for the M/E at point N qM/E,N [m3/h] 62,781 Refer to (9)
required air flow for d/gs at %ND/G qD/G,N [m3/h] 4,356 Refer to (10)
required air flow for oil-fired boiler qb [m3/h] 24,372 Refer to (11)
required air flow for combustion qc [m3/h] 91,509 Refer to (6)
required air flow for heat evacuation qh [m3/h] 53,424 Refer to (7)
total air flow required at current vessel operating profile
At seagoing condition (M/E and D/G) 100,700 Reference to ISO 8861:1998
and (5)At port/anchorage/loading condition qtot,N [m3/h] 6,500
At unloading condition (d/g and boiler) 43,100
Data-DrivenProcess
Controller
VFD
E/RArea
OutsideArea
DP
Amb.Temp
Real-Time MeasurementsParameter Set PointsCommunication
HM
I
Differential Pressure
Fan Numbers1–4
Data Transfer
Pdiff,E /R*
Pdiff,E /RTamb,E/R
*
Tamb,E/RNfan
ωref*
PE/RPout
The proposed data-driven controller interface with the process.3
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( ) ,
T T
T
12
7 12
C
C C
, ,
.
max max
maxS W
/amb E R amb outo
o o
= +
= + +
(26)
where for maximum T .S W considered equal to 32 °c, then the maximum e/r ambient temperature for a design of e/r components is 51 °c, or 55 °c for extreme conditions.
however, since air ventilation ducts for a normal air intake system are placed near the turbochargers, air inlet temperature to the turbocharger will be very close to T ,amb out, as in
take( ) ,T Tturbocharger in 3 C, ,max max
/amb E R amb outo= + (27)
where maximum air temperature near the turbocharger intake should be 42 °c, or 45 °c for extreme conditions. this temperature is associated with the temperature of the m/e scavenge air chamber T ,air scav, which depends on scavenge fresh water cooler temperature T .WF , which is
equal to 36 °c for T .S W = 32 °c. since the standard marine scavenge air cooler is specified with a maximum 12 °c difference between T .WF and T ,air scav at 100% of the mcr, the maximum scavenge air temperature must be 36 °c + 12 °c = 48 °c, allowing a 7 °c margin to scavenge air cooler high temperature alarm (set at 55 °c), after which the m/e load must be reduced. to satisfy this limit and at the same time avoid condensation in the scavenge air chamber, the controller temperature set point
The design of the proposed data-driven control topology and VFD interface.4
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is selected at 36 °c, based on measurements performed via an ambient sensor placed at the turbocharger level. a low speed limit it,limlow~ equal to 40% of now~ is also
established to avoid motor overheating [27].during a port stay, the m/e is preheated to prevent
temperature variations and thermal expansions in its struc-ture that may lead to leakages. in this mode, the controller also tries to maintain constant T ,amb E/R , while the m/e
jacket cooling water outlet temperature is kept high (max-imum 75–80 °c) and, before startup, is increased to at least 50 °c through a built-in heat preheater. additional-ly, during unloading, T ,amb E/R is maintained constant to assist with heat evacuation.
When a ship occasionally operates under arctic condi-tions with low m/e turbocharger air intake tempera-tures, say 0 °c, the air density t will rise, causing scavenge air pressure P ,M/E scavenge to be too high, affect-ing compression pressure and maximum firing pressure. to prevent such excessive pressures under low ambient temperature conditions, the turbocharger air inlet tem-perature should be kept as high as possible. further-more, in this case, T .S W will be also low, so T .WF should be kept as low as possible, and the m/e load should be reduced. since the m/e will work at a very low load in
these ambient outside temperatures, both combustion and convection heat air-flow demands from the e/r ven-tilation system will be minimized. in this condition, since both the e/r differential pressure and the e/r ambient temperature pid speed reference control out-puts, ,P diff~
) and ,,T E/R~) respectively, will be constantly
very low, close to the minimum allowable speed limit min~ , the controller will constantly operate the mini-
mum number of fans (one fan) at a very low speed, close to the minimum limit of min~ = 50%, according to table 3 and (23).
the controller dynamic parameters, set points, and minimum limits are defined in table 3. a control override function is also included to run the fans at 95% of nom~ in case the control unit or sensor fault is detected. to illus-trate the different control algorithms implemented in the proposed controller, the first cycle of the control system evolution is presented in figure 5 as a sequential function chart (sfc). it is noted that, if the mentioned safety checks fail, the standby e/r fan vfd unit will start.
Design of VFD Topologythe applied vfd topology for each fan motor consists of a six-pulse diode rectifier and a standard two-level, three-phase voltage source inverter with a 3% direct current (dc) choke included in the dc-link. With reference to the current flowing in the alternating current (ac) supply i ,i a and the line-to-line supply voltage ,vLL the required choke inductance L is calculated according to the desired voltage drop across the reactor v ,L ch at fundamental fre-quency f for an equivalent ac line reactance as in
[ ]% ( / )
,L Hf IV
f IV
2 100 23
,
,
,
,
i a
L ch
i a
L ch LL
$ $ $$
r r
oD= =uu
uu
(28)
where parameter %v ,L ch3 is defined in (29) as a percent-age ratio of the desired voltage drop across the reactor to the drive input phase voltage at fundamental frequency:
%/
.VV
1003
,,
LLL
Lch
ch$oD =u
u (29)
the harmonic filters are selected so that the harmonic limits adopted by marine classification societies are satis-fied [28].
Implementation of the Proposed Control Topologythe designed data-driven process control topology for the optimization of e/r ventilation systems is imple-mented in the tanker vessel under study as a retrofit installation to control the speed of four 18.5-kW, 440-v, 33.6-a, 60-hz e/r ventilation fan motors, according to the proposed data-driven control algorithm. the arrangement of the process controller as well as its interface with sensors, vfd, and hmi units are pre-sented in figure 6. the control unit includes a modicon m340 plc along with analog and digital input–out-put cards and a communication module, interfacing with an atv61hd22n4 vfd unit. to perform pro-cess monitoring, the working plant is implemented in a magelis hmistU855 graphic display unit, using vijeo designer hmi configuration software. the ana-log input signals are received by 4–20 ma outputs
TABLE 3. THE DEFINITIONS OF CONTROLLER PARAMETERS, SET POINTS, AND MINIMUM LIMITS.
Nfan
(in use)Defined Set Points at Each Condition
lower Fan Speed Limit
Applied Process Constraints
One fan (cold area)
P 15Pa,diff E/R =)
T 36 C,amb E/R c=)
%50min~ = .P 0 7>,M/E scavenge bar (for turbocharger normal operation)
Two fans (port)
P 25Pa,diff E/R =)
T 36 C,amb E/R c=)
%07min~ = T 48 C<,air scav c
Three fans (seagoing slow)
P 35Pa,diff E/R =)
T 36 C,amb E/R c=)
%06min~ = .P 0 1>,diff E/R mbar it %40,limlow~ = (to avoid motor overheating)
Four fans (seagoing high)
P 45Pa,diff E/R =)
T 36 C,amb E/R c=)
%05min~ =
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from temperature and differential pressure transmitters used to provide real-time feedback signals to the con-troller. the applied differential pressure sensor has a range of –2.5 to +2.5 mbar to be able to detect both overpressure and underpressure using a diaphragm that expands or contracts with pressure changes. a pt100 temperature sensor of 0–60 °c is used to record e/r ambient temperature. analog 4–20 ma and 0–10 v controller output signals are used to provide speed ref-erence to the vfd embedded microcontroller module. modbus over transmission control protocol/internet protocol (tcp/ip) is used to synchronize vfd opera-tion with the plc, while ethernet over tcp/ip is used for hmi communication with the plc control unit. digital input signals inform the controller about sys-tem status, including overload, phase failure, tempera-ture rising, and vfd failure conditions to generate relevant alarms.
the hardware configuration of the proposed control unit installed on the tanker vessel is presented in figure 7(a), while the arrangement of vfd topology for each fan is presented in (b). the applied vfd topology allows all four fans to be operated for either supply or exhaust if needed. during seagoing mode at slow steaming operation with 9,282 kW (65% of the mcr), usually three out of four fans are in service. as shown in figure 8(a), hmi depicts the operating status of the working plant as a result of the proposed data-driven controller actions. it is shown that the controller oper-ates three fans at 67.5% of their nominal speed, namely 40.5 hz, keeping the e/r overpressure P ,diff E/R equal to 0.33 mbar or 33 pa, while, at the same time, the e/r ambient temperature is maintained at 32.1 °c. as shown in figure 8(b), each fan consumes 5.9 kW instead of nominal 18.5 kW, resulting in 37.8-kW sav-ings in total.
Start
1Data-DrivenControl Start
1 7PLC Control Unitand Sensors Ok
PLC Control Unitor Sensors Fault
2HMIOperation
8 Set OverrideSpeed 95%
9
ActivateVFD
“Auto”Operational Mode
Start StandbyFan
SafetyCheck Ok
VFD Status Ok
SafetyCheck Faults
3
4
5 6
10
2
3
4
5
8
9
10
7Limit Check
6
Nfan
Pdiff_E/R*
Pdiff_E/R
ωP_diff*
ωmin
ω T_DP*
ω T_E/R*
Tamb_E/R*
Tamb_E/R
ωauto*
ω fault*
ωref*
PM/E_scav ≥ 0.7 bar
ω ref ≥ 40%
Tair_scav < 48 °C
An SFC showing the first cycle of the proposed data-driven control system evolution.5
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Experimental Results and Discussionthe dynamic performance of the proposed controller in different vessel operating conditions, as well as its impact on power balance optimization and improvement in d/g fuel consumption, will be evaluated in this sec-tion. for performance validation, experimental results are obtained during a typical voyage after the retrofit instal-lation of the proposed topology.
Dynamics and Performance of the Proposed Controllerthe dynamics of the proposed controller regarding pro-cess parameter variation are examined in real time ini-tially during a seagoing period and are presented in
figure 9. during typical vessel seagoing operation in slow steaming condition, where the m/e is working at
nN = 88.6 r/min producing a load of PN = 8,568kW, three out of four e/r ventilation fans are working with the proposed data-driven control. in figure 9, it is shown that since Nfan= 3, the controller automatically
applies the differential pressure set point of P ,diff E/R)
=
35 pa and the e/r temperature set point of T ,amb E/R)
=
36 °c, according to table 3. therefore, it is shown that, at this stage, the differential pressure pid control loop is tuned to provide a speed set point ,P diff~
) , which is equal to 65.23% of now~ to satisfy the demands of the m/e turbocharger in terms of combustion air flow and to keep the air pressure in the m/e scavenge
PLCControl
DP
TT
Pt1
00 Numbers 1–4 VFD
Differential Pressure Transmitter, E/R PressureDifference –2.5… +2.5 mbar, 4–20 mA
Temp. Transmitter, E/R Temperature,Range 0...60 °C Pt100, 4–20 mA
Auto SpeedReference
Modbus Over TCP/IPAI1
AI2AO1
AO2Ethernet Over
TCP/IPRJ45
RJ45
RJ45
4–20 mA
4–20 mA4–20 mA
RJ45
DO1–DO824 Vdc
DI1–DI8
DI1–DI8
DO1–DO8
24 VdcHMI
AI2
AI1
0–10 V
Manual Speed Reference
The hardware implementation layout of the proposed control topology.6
PLC CPU Module PLC DIM, AIM,AOM Cards
ProgrammableController
VFD Unit
dc-Choke
(a) (b)
Modbus Hub Communication Module
The implementation of the proposed data-driven control topology in the tanker vessel under study. (a) The arrangement of the central PLC control and communication units and (b) the layout of the VFD topology installed for each fan.
7
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air chamber above 0.7 bar. at the same time, the e/r temperature pid control loop provides a much lower speed set point of ,T E/R~
) , which is equal to 20% of nom~ , since the actual e/r ambient temperature is well below the command value. however, this set point is ignored because it violates it,limlow~ = 40%, set to protect the motor from overheating.
in any case, the final speed reference provided to each fan motor is ref~
) = auto~
) = 65.2%, which is the maxi-
mum reference between ,P diff~) and ,T E/R~
) , according to (23) and complies with the minimum limit of min~ =
55%, as defined in table 3. this set point is the same for all the fans in service, leading to a power saving of 35 kW
compared to the previous mode. the variation of the
required combustion air flow by the m/e and d/gs due to the vessel operating at different speeds will change the actual ambient e/r pressure, which will trigger the
controller to change the speed of the fans according to the pressure and temperature feedback signals and the num-ber of fans in service. to this end in figure 9, it is shown that controller dynamics are well established, since the variation of actual differential pressure from 50 to 37.5 pa leads to reduction in fan motor speed reference from 65% to 52%, while the output speed reference of the controller displays excellent transient characteristics, including minimum overshoot and rise time. it is noted that the
(a) (b)
The process monitoring functions of the data-driven controller HMI screen. (a) The implementation of the working plant operating condition and (b) the monitoring of the process parameters and motor power consumption.
The experimental results: the dynamics of the proposed data-driven controller at seagoing operation with the M/E working at 60% of the MCR (nN = 88 r/min, PN = 8,240 kW), where three out of four of the E/R ventilation fans are in service (PID tuning parameters: KP,diff =10, KI,diff =10, KD,diff =2.5; and KP,T = 10, KI,T = 2.5, KD,T = 10).
9
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controller sensitivity to sudden variations in ambient pressure is regulated to enhance the stability of the fan motor revolutions per minute command.
in figure 10, experimental results are provided during the vessel cargo-unloading operation, where the m/e is not in use, but the oil-fired boiler works at its designed capac-ity, i.e., qb = 25,000 m3/h with three fans also in service. at this stage, although P ,diff E/R
) and T ,amb E/R) are kept the
same as before at 35 pa and 36 °c, respectively, the motor speed is regulated by the e/r temperature pid control loop output at ref~
) = ,T E/R~
) =87.52%, according to (23),
since the need to remove the convection heat generated by the oil-fired boiler operation is higher than the combus-tion air-flow demand. since ,T E/R~
) is higher than ,P diff~
) due to the m/e being idle, P ,diff E/R is increased.
for the first e/r fan, typical operation at 45 hz, motor side voltage ( ),v ,m LL and current ( )i ,am waveforms
are shown in figure 11(a), while the corresponding grid side voltage ( )v ,i LL and current ( )i ,i a waveforms are pre-sented in figure 11(b). While the motor impedance acts as a filter that assists with the elimination of the high-frequency harmonic components on the motor side, a 600-μh choke is used in the dc-link of each drive to reduce current harmonic distortion on the grid side, according to (28)–(29). the harmonic spectrum of cur-rent waveform presented in figure 11(c) shows that har-monic components include fifth, seventh, 11th, 13th, and 17th harmonic order. in figure 11(d), it is shown that thdv% < 5%, and the amplitude of individual voltage harmonics at pcc is less than 3% with reference to the fundamental frequency, both of which comply with marine harmonic standards [28].
Evaluation of Power Balance Improvementthe variation in the active power consumption of the first e/r ventilation fan driven by the proposed control topology is recorded and presented in figure 12 for a typical ten-day, seagoing voyage in comparison to its
original power consumption before the proposed system retrofit installation. the power consumption waveform is identical for all fans in service at that given moment. for the usual case, where all three or four e/r fans are in service to achieve uniform air distribution in the e/r, it is shown that the control system responds well to the variable e/r component combustion and heat emission air-flow requirements. When the m/e is working at 75% of the mcr (PN = 10,710 kW) with four fans in service, the controller regulates their speed at 43 hz to compensate for additional combustion air-flow requirements and maintain 45 pa overpressure in the e/r, resulting in about 7-kW power consumption for each fan. this leads to about 46-kW total power saving in comparison to the previous condition.
however, the effectiveness of the controller is demon-strated at a slow steaming period, where the m/e load PN is equal to 55% or 7,854 kW and its speed nN is equal to 82% or 86 r/min. during this period, it is shown that if four fans are in service, each fan operates near minimum 30 hz with individual fan power consumption of 2.5 kW, which is very close to the theoretically calculat-ed values from the air balance case study performed in the section “e/r air Balance calculation for the ship Under study.” the total power-saving margin in this case is 64 kW, considering that nominal fan power consump-tion was 18.5 kW. if three fans are used, P ,diff E/R
) is auto-matically adjusted at 35 pa, and it is shown that the controller raises the speed of each fan at 43 hz, resulting in 21-kW total power consumption or 35-kW total power saving, respectively.
Fuel Efficiency and Economic Benefits from Control Systemin this section, the impact on d/g power and fuel con-sumption is examined as well as on emission factors (co2, so2, and nox) through a comparative study. con-sidering that heavy fuel oil (hfo) is used as standard
The experimental results: the dynamics of the proposed data-driven controller at the cargo-unloading condition with the boiler working at nominal capacity (qb = 25,000 m3/h), where three out of four of the E/R ventilation fans are in service (PID tuning parameters: KP,diff =10, KI,diff =10, KD,diff =2.5; and KP,T = 10, KI,T = 2.5, KD,T = 10).
10
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d/g fuel, the d/g fuel oil consumption (focd/g) induced from e/r fan power consumption is based on specific fuel oil consumption (sfocd/g) and running time (rt) in hours as
.PFOC RT SFOC_/ /D G HFO fan D G# #= (30)
for the d/gs used on board, a sfocd/g of 192.3 g/kWh is considered if one d/g is running alone at 75% of its capacity, while a sfocd/g of 210 g/kWh is con-sidered if two d/gs are running in parallel at 40% of their capacities.
the corresponding d/g emission factors due to hfo fuel consumption are calculated based on
[ ] [ ]
[ ] .
tn tn C
tn 3 17
CO FOC
FOC
2 /
/
D G FO
D G
#
#
=
=
(31)
[ ] [ ] % .tn tn S 0 02SO FOC2 /D G # #= (32)
=[ ] [ ] . ,tn tn 0 079NO FOCx /D G # (33)
where Cfo is the conversion coefficient of fuel oil and car-bon emission for marine heavy duty diesel engines and S% is the sulfur content included in fuel oil [29].
the comparative study is based on the operating pro-file of the vessel, where experimental results were obtained from the proposed controller application. the comparative results over the original e/r ventilation fan operation with a direct online (dol) starting method for a year of operation are shown in table 4 and in fig-ure 13, including vessel seagoing, port in/out, anchor-age/loading, and unloading periods. the energy savings in kilowatthours are calculated based on the difference in e/r fan power consumption between dol and pro-posed data-driven vfd control for the specific run-ning time at each period. the d/g fuel savings in tons of hfo are then calculated based on (30), while the cal-culation regarding the reduction in emission factors is based on (31)–(33), considering a 2% sulfur percentage
(c) (d)(b)(a)
The typical waveforms of the proposed VFD topology for the first E/R ventilation fan working at 45 Hz. (a) The motor side volt-age (vm,LL) and current (im,a) waveforms, (b) the grid side supply voltage (vi,LL) and current (ii,a) waveforms, (c) the harmonic spectrum of ii,a, and (d) the harmonic of vi,LL, as measured at PCC.
11
E/R Fan Power ConsumptionWithout Data-Driven Control Topology
Fan PowerConsumption at 60 Hz
Three Fans (Working at 43 Hz)
Four Fans(at 30 Hz)
Four Fans(at 43 Hz)
M/E Working in Slow Steaming Condition(nN = 86–88.6 r/m, nN% = 81.9–84.4%PN = 7,854–8, 568 kW,PN% = 55–60%)
Thursday 9 May 2013, 11:52:48 – Monday 20 May 2013, 04:51:28
Min/Max Total Active Power (Fundamental + Harmonics) (Cycle by Cycle), ELSPEC11@ELSPEC-NTUA
A comparison of the E/R fan power consumption between the data-driven process control topology and the previous operating condition for a typical voyage.
12
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in hfo. Based on 2013–2014 average hfo prices of Us$650 per ton, it is shown that total annual benefits include a 55-ton/year reduction in hfo consumption for the vessel, meaning about Us$36,000 in economic benefits. considering that co2 emissions are reduced by 174.35 tons/year, there are also significant
environmental benefits.to evaluate the influence of
power savings with respect to the total d/g load in the vessel’s power system, figure 14 compares the total power system load before and after the proposed topology retrofit installation during vessel seagoing
(nominal speed and slow-steaming), loading/anchorage, and unloading periods. although it appears that the highest power-saving percent-age occurs during the loading/anchorage mode, where the load is minimum, the effects on the power balance are more evident during the seagoing period at a slow steaming
condition, which occupies 46% of the total vessel’s oper-ating time. to further elaborate on this, in the autono-mous power system of marine vessels, the maximum permitted load serviced by one d/g is around 80% of
TABLE 4. THE CALCULATIONS OF THE FUEL CONSUMPTION, EMISSION FACTORS, AND ENERGY-SAVING BENEFITS FROM THE PROPOSED DATA-DRIVEN CONTROL SCHEME ACCORDING TO THE VESSEL OPERATING CONDITIONS (HFO PRICE = US$650 PER TON).
vessel operating Mode and Equipment used
Normally SeagoingPort/Loading Anchorage Unloading
M/E in Use (PN = 85%) One D/G in Use
M/E in Use (PN = 60%) One D/G in Use
M/E Not in Use (PN = 0%) One D/G in Use
M/E Not in Use (PN = 0%), Two D/Gs and Boiler in Use
Operating time 350 h 4,030 h 3,854 h 526 h
(4%) (46%) (44%) (6%)
Nfan (in use) 4 3 2 3
E/R ventilation fans total power (DOL)
74 kw (four fans)
55.5 kw (three fans)
37 kw (two fans)
55.5 kw (three fans)
E/R ventilation fans total power (proposed control scheme)
55-ton/yearHFo reduction, us$35,700 per year economic benefit from fuel costs
total emission factors reduction 174.35-ton co2 savings/year, 2.2-ton so2 savings/year, 4.3-ton nox savings/year
A ship’s annual savings from the proposed data-driven control system.
0
20
40
60
80
100
120
E/R
Fan
Tot
alP
ower
Con
sum
ptio
n (k
W)
Year Savings4,900 kWh
1.02 tons HFO
Year Savings141,050 kWh
29.30 tons HFO
Year Savings96,350 kWh
20.01 tons HFO
Year Savings20,770 kWh
4.71 tons HFO
7460 56
37
56
21 12 16
Seagoing(Nominal Speed)
Four Fans in Use
Seagoing(Slow Steaming)
Loading/Anchorage
Three Fans in Use Two Fans in Use
Unloading
Three Fans in Use
D.O.L Starting Mode Data-Driven Variable-Frequency Control
13
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its mcr (780 kW in the vessel under study) before another d/g is started, due to restrictions concerning single point failure, maximum transient load, and avail-able power reserves. during the slow steaming seagoing operation, usually m/e auxiliary blowers work to sup-port the turbocharger, while the auxiliary boiler is also in use along with its associated consumers (forced draft fans and feed water pumps), because at low m/e load the exhaust gas boiler is not sufficient to sustain adequate steam flow for the e/r and cargo heating purposes, lead-ing to an average continuous load of 624 kW. since this load is close to 80% limit, a second d/g must often run in parallel at a low load with a high sfoc, leading to inefficient operation. table 4 and figure 14 show that the created continuous power-saving margin of 35 kW from the proposed system at this period results in a total 589-kW d/g load, somewhat below the 80% limit, meaning that only one d/g can often be used to service
load demand, leading to power balance improvement.
the cost breakdown of the pro-posed data-driven control topology is presented in table 5. the relevant costs, based on 2013 component catalog prices, include the e/r fan
vfd starter assembly with associat-ed components, the controller topol-ogy based on plc use, the hmi screen for remote control, and the sensors used to obtain the feedback signals. the cost-benefit analysis shows that the payback period is under a year, justifying the recom-mendation for the application of the proposed energy-efficiency upgrade. it has to be taken into account, however, that economic and fuel benefits from the proposed imple-mentation strongly depend on sever-al key parameters involving the installed fan capacity, the operating
profile of the vessel, and fuel prices; therefore, an analyti-cal case study is recommended before application.
Conclusionsin this article, a data-driven process control topology was designed for energy-efficient, adjustable-speed fan operation in the e/r ventilation system of marine vessels. the control-ler parameter adjustment was based on real-time process measurements, which were used as feedback signals to tune the controller. the controller was used to adjust the speed of four e/r ventilation fan motors according to variable com-bustion air-flow requirements and the vessel’s operating pro-file to increase energy efficiency. the proposed controller was implemented using a commercial plc, while an hmi was developed to monitor related process parameters. air balance calculations were initially performed to examine power-sav-ing potential and critical aspects of the process. the experi-mental verification of the proposed control system retrofit
2,1
5,6
7,1
3,9
Sav
ing
% o
fTo
tal P
ower
Sys
tem
Loa
d
D/G Optimization
0
2
4
6
8
10
12
Psys,old = 670 kW
Psys,vfd = 656 kW
Psys,old = 350 kW
Psys,vfd = 325 kWPsys,old = 624 kW
Psys,vfd = 589 kW Psys,old = 1,014 kW
Psys,vfd = 974.5 kW
Seagoing(Nominal Speed)
350 hFour Fans in Use
Seagoing(Slow Steaming)
Loading/Anchorage
4,030 hThree Fans in Use
3,854 hTwo Fans in Use
Unloading
526 hThree Fans in Use
Power Saving as Percentage (%) of Total Power System Load
A comparison of the vessel’s total power system load before and after the proposed energy efficiency topology retrofit installation.
14
TABLE 5. THE COST BREAKDOWN OF THE PROPOSED CONTROL SYSTEM AND PAYBACK PERIOD.
controller topology (including plc with cpu, digital/analog I/o and communication modules)
US$4,800 total
HMI screen for remote control US$1,150 total
Sensors and transmitters US$450 total
Cabling and peripheral material for installation US$1,600 total
Total cost for onboard retrofit application US$30,000 total
Average annual fuel savings (2013–2014 prices) US$35,700 total
Investment payback period based on savings Ten months
This article has been accepted for inclusion in a future issue of this journal. Content is final as presented, with the exception of pagination.
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installation on board an existing tanker vessel verified that system dynamic responses were well established, achieving significant energy and fuel savings, approaching 260 mWh and 55 tons of hfo reduction in consumption per year, respectively. comparison results over the original e/r venti-lation fan operation method showed significant power bal-ance improvement and emission factor reduction, verifying that the marine industry can greatly benefit from this ener-gy-efficiency upgrade.
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[25] Shipbuilding-Engine Room Ventilation in Diesel-Engined Ships—Design Requirements and Basis of Calculations, iso standard 8861, 1998.
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Spyridon V. Giannoutsos ([email protected]) is with the National Technical University of Athens, Greece, and Thena-maris Ships Management, Athens, Greece. Stefanos N. Manias is with the National Technical University of Athens, Greece. Giannoutsos is a Student Member of the IEEE. Manias is a Fel-low of the IEEE. This article first appeared as “A Data-Driven Process Controller for Energy Efficient Variable-Speed Fan Oper-ation in Engine Room Ventilation System of Marine Vessels” at the 2014 IEEE IAS Annual Meeting.
Abstract—Nowadays marine vessels operate their Main
Engine (M/E) at low speeds to minimize fuel consumption. This
practice leads to significant load balance variation in vessel’s
autonomous power system compared to sea trials, frequently
requiring an additional Diesel Generator (D/G) to work in
parallel at low load to satisfy the limits set by power management
system. To improve D/G operating profile and Specific Fuel Oil
Consumption (SFOC), a data-driven control topology is
proposed, involving speed adjustment of motors related to engine
room ventilation and central cooling water processes that exhibit
varying flow requirements according to ship’s operating profile.
A power saving margin that allows use of only one D/G at sea-
going is created, optimizing D/G fuel consumption for M/E slow
steaming operation. The proposed system is applied as retrofit
installation in a typical tanker vessel’s power system, where
experimental results are compared to the previous operating
condition to show the impact of D/G optimized operation on fuel
consumption. A method is proposed to calculate economic and
environmental benefits depending on type of fuel used, D/G
SFOC and vessel’s trading pattern. Results show that vessel’s
power system annual energy consumption and D/G fuel
consumption are significantly reduced by 18.3% and 20.3%
respectively.
Keywords— Marine vehicle power systems; Industrial power
system control; Energy management; Power conversion; Process
control; Variable speed drives
I. INTRODUCTION
Marine power systems undergo structural changes with new control strategies utilizing distribution automation and network reconfiguration to provide greater flexibility, reliability and reduce fuel consumption [1]-[2]. Optimal Diesel Generator (D/G) allocation control is achieved by the vessel’s Power Management System (PMS) to satisfy safe continuous load, react to transient load step variations as well as maintain sufficient power reserve to prevent blackout conditions [3]-[5]. On operational level, a method for improved demand-side management and power generation scheduling is proposed in [6], while real-time power management techniques that support system critical operations in the event of dynamic load change based on switch status of loads are discussed in [7]-[9]. The growing integration of power electronic converters in several
parts of the ship’s power system including propulsion, power distribution and auxiliaries also affects power balance with the intention to reduce fuel consumption and greenhouse gas (GHG) emissions [10]. To improve efficiency, optimal power management methods are proposed in [11]-[12] for marine power systems integrating energy storage systems, while options for retrofit application of capacitor banks for power factor correction purposes are discussed in [13].
Nowadays, however, the commonly used slow steaming practice, involving vessel’s Main Engine (M/E) operation at lower speeds [14], leads to significant load variation in vessel’s autonomous power system at sea-going period. In contradiction to optimally pre-configured D/G operation based on calculated continuous and intermittent load levels, also tested during sea trials, the load variation due to slow steaming often leads to an additional D/G working in parallel to satisfy the PMS limits for adequate online power reserve. In that case, parallel D/G low load operation increases individual Specific Fuel Oil Consumption (SFOC), subsequently resulting in increased fuel consumption and GHG emissions especially during sea-going.
As a direct consequence, slow steaming leads to derated flow rate requirements in processes relevant to M/E and auxiliary machinery operation due to their variable water cooling, combustion air and heat dissipation needs [15]. To increase energy efficiency, adaptive speed control of motors associated with these processes according to varying process requirements has been proposed in [16]-[17]. However, the dynamic power saving margin created by the variable frequency control of motors, which would otherwise operate constantly at nominal power, could significantly be used to improve vessel’s power balance especially during sea-going.
To exploit this power saving margin in order to optimize D/G operating profile and fuel consumption during sea-going operation at reduced M/E speed while simultaneously satisfy PMS limits, a data-driven control topology is proposed in this work, involving Variable Frequency Drive (VFD) application to motors associated with Engine Room (E/R) ventilation and central cooling water system processes. By compensating the vessel’s load variation due to its slow-steaming operating profile with the created power saving margin from the proposed
Fig. 1. Variation of Specific Fuel Oil Consumption (SFOC) in g/kWh with
the load of D/G as a percentage of its nominal shaft mechanical power
(measurements taken from an L23/30H D/G with nominal power of 780kW)
topology, the constant use of only one D/G is achieved during sea-going operation, significantly optimizing power balance and D/G SFOC levels. The proposed topology is applied as retrofit installation in the power system of a typical tanker vessel, where the obtained experimental results are compared directly to the previous operating condition to show the impact of D/G optimized operation on fuel consumption. Through a method that is introduced to calculate economic and environmental benefits depending on type of fuel used, SFOC variation and vessel’s trading pattern, it is shown that the improved power balance significantly reduces D/G fuel consumption compared to the original operation, verifying that the marine industry can be greatly benefited from this energy efficiency upgrade.
II. LOAD BALANCE AND D/G OPERATION AT SEA-GOING
A. D/G operating profile and relation to fuel consumption
In marine autonomous power systems, the number of online D/G units, k, and their loading level, PL,gi, is configured by PMS based on load dependent start tables and on
restrictions concerning single point failure, maximum transient load and available power reserve. In case of D/G failure the load must be transferred to remaining D/Gs online, while the transient frequency deviation must be limited to ±10% according to class society rules [18], in order to avoid tripping of ACBs and prevent blackout. The transient load step in the system, ΔPtran(k,Nf), when Nf units are tripped for k units online is defined as in Eq.1:
fN
f
fgfLftranL NkPNkP1
,, ),(),( (1)
where PL,gf is the pre-fault system load. Since resistance to single failure is required, the limits are calculated for the case which the unit with the highest loading fails as in Eq.2:
],1[)},({max)( ,, kiforkPkP giLigfL (2)
The corresponding transient load per D/G is determined to be the sum of D/G load and transient load step, as in Eq.3:
),(),( ,,,,, fgitranLgiLfgitranL NkPPNkP (3)
The blackout can be prevented as long as the maximum transient load step is lower or equal to the permitted, leading to maximum continuous safe limit (blackout limit), as expressed in Eq.4 for each D/G remaining online:
][)],([
)(),(),(
,,max,,,,
max,
max
,,,,
max
,,
giLgirgfgitranLgiL
ggitranLfgitranLfgicontL
PPaNkPP
aPNkPNkP
(4)
where Pr,gi is D/G power rating and amax,g is the maximum permitted engine load constant, typically between 1.1 to 1.15. With respect to k D/G units remaining online, the next unit must start before the load reaches maximum continuous safe limit. Thus, the safe operating region for starting D/G, PL,gi(k), is limited to the minimum value between maximum safe load and rated D/G load, as shown in Eq.5:
girfgicontLgiL PNkPkP ,
max
,,, ),,(min)( (5)
From the above, it is shown that the system must always have a sufficient power reserve, Pr,av, or available power to its full online capacity to prevent blackout as shown in Eq. 6:
k
i
gtotLgirgtotLgravr PPkPkPkP1
,,,,, )()()( (6)
where minimum available power, based on generator failure cases and units online is determined according to Eq.7:
k
i
giL
k
i
girfgicontLfavL PPNkPNkP1
,
1
,
max
,,
min
, }),,(min{),( (7)
Another restriction regarding PL,gi for online units is the
recommended minimum engine load, as described in Eq.8:
30.015.0,)( min,,min,, ggirggiL aPakP (8)
The PMS also performs equal load sharing between D/Gs according to classification society regulations, to avoid reverse power or load limiting conditions.
Based on the above, the usual maximum permitted D/G load for a marine tanker vessel based on pre-configured start-stop tables for obtaining lowest possible fuel consumption is 80% of Pr,gi, while minimum load is 20% of Pr,gi. However, measurements show that D/G SFOC in g/kWh greatly depends on loading percentage, as it is a convex curve with minimum values near 70% of Pr,gi. For the vessel under study, where 3 sets of 780kW 6L23/30H D/Gs are installed [19], the SFOC
Fig. 2. Load balance variation during sea-going operation as percentage of
individual D/G nominal load with M/E load and speed for a typical marine
vessel (measurements taken from a 105,000DWT Aframax tanker vessel)
curve for Marine Diesel Oil (MDO) use is presented in Fig.1, where it is shown that SFOC dramatically increases if D/G load is less than 55% of Pr,gi. Therefore, if D/Gs have to work in parallel at low loads, this would have a considerable impact on D/G fuel consumption, as per Eq.9:
]/[][][][ , kWhgSFOChRTkWPtnFOC gimech (9)
where Pmech,gi is the shaft mechanical power of D/G, which can be calculated based on measured output power, PL,gi, generator rated power, Pr,gi, and generator copper loses, Pcop, core losses, Pcore, mechanical losses, Pm, and stray losses, Pstr, as in Eq.10:
]%1%4%33[
][
,,,
2
,
,,,
girgirgiLsLgiL
strmcorecopgiLlossgiLgimech
PPPRIP
PPPPPPPP
(10)
where IL is the current flowing in each armature winding phase
and Rs is the armature winding resistance.
Consequently, the corresponding GHG emissions due to D/G FOC can be calculated as shown in Eq. 11-13:
17.3][][][ //2 tnFOCCtnFOCtnCO GDFOGD (11)
02.0%][][ /2 StnFOCtnSO GD (12)
079.0][][ / tnFOCtnNO GDx (13)
where CFO is the conversion coefficient of fuel oil and carbon emission for marine heavy duty diesel engines and S% is the sulfur content included in fuel oil [20].
B. Vessel’s D/G load variation at slow-steaming condition
For a typical marine vessel, power system load, PL,g, during sea-going operation greatly varies with M/E load and speed, i.e
TABLE I
E/R MACHINERY DETAILS FOR THE VESSEL UNDER STUDY
with the working percentage of M/E Maximum Capacity Ratio (MCR), PM/E,MCR, due to variation of load requirements for related M/E operation processes. For the vessel under study, PL,g is presented in Fig.2 as varying percentage of Pr,gi with M/E working load, PM/E,N. It is shown that in high M/E load area range (63% to 80% of MCR) only one D/G is needed, since PL,g is below 80% of Pr,gi, while in very low M/E load range (30% to 45% of MCR), the use of a second D/G is imperative since PL,g is around 115% of PM/E,MCR, significantly above 80% of Pr,gi. This occurs because at low M/E load, economizer cannot sustain adequate steam flow and 5.0bar steam pressure for E/R and cargo heating purposes, leading to continuous use of auxiliary boiler, which involves additional consumers (forced draft fans, feed water pumps). Additionally at low load, M/E turbocharger cannot achieve adequate overpressure in M/E scavenge air chamber, so both M/E auxiliary blower motors must continuously run to keep scavenge air pressure more than 0.7bar. In this area, considerable quantity of fuel saved from M/E working at lower speed is consumed by D/Gs due to increased loading.
As M/E load increases and enters the area 45% to 55% of MCR, the use of aux. boiler and M/E aux. blowers becomes intermittent with frequent start/stops, so this area is unsuitable for constant sea-going operation. Therefore, ship’s M/E usually operates in the range 55% to 65% of MCR to achieve optimal savings. As shown in Fig. 2, however, PL,g in this range is close to 80% of Pr,gi, often leading to another D/G running in parallel. In that case, however, each D/G would run at low load, close to 40% of Pr,gi, resulting in 10% SFOC increase, thus having significant impact on FOC. The topology proposed in this work aims to match actual E/R consumer needs to reduce D/G load at sea-going condition so that only one D/G is used. The impact of proposed system to fuel consumption is evaluated through its retrofit in a real Aframax tanker vessel, whose E/R machinery details are shown in Table I.
III. PROPOSED VFD-BASED PROCESS CONTROL TOPOLOGY
A. VFD topology retrofit to existing vessel’s power system
The 440V, 60Hz loads in the power system of a typical marine vessel, as presented in Fig.3, are fed through group starter distribution panels from main switchboard or directly from emergency switchboard, while motors are connected either D.O.L or Star-Delta. To achieve the desired power saving margin especially during sea-going period, the proposed topology is applied in E/R ventilation and central cooling water processes, which exhibit considerable variation
Fig. 3. Integration of the proposed VFD-based process control topology in
the typical power system of the tanker vessel under study
TABLE II
VARIATION OF PROCESS FLOW REQUIREMENTS AT SLOW-STEAMING
of flow requirements with M/E load and speed. The topology involves VFD application to all four E/R ventilation fan motors and to one sea water (S.W) cooling pump motor, all of which work continuously throughout the vessel’s operating profile.
The speed control of E/R ventilation fans is performed according to feedback from real-time measurements of differential ambient pressure between E/R and outside, Pdiff,E/R, and ambient temperature, Tamb,E/R, in order to meet varying combustion and heat evacuation air flow requirements, q
*fans, in
relation to M/E speed for major E/R consumers, including M/E, D/Gs and oil fired boilers. Similarly, speed control of S.W cooling pump is performed based on central cooling water system parameter variation, involving feedback from real-time measurements of S.W pressure, PS.W, fresh water (F.W) temperature, TF.W, and S.W temperature, TS.W, in order to satisfy the heat dissipation requirements, q
*pump, of M/E jacket,
scavenge and lubrication coolers according to varying M/E speed and load. The appropriate speed references for fan and pump motors, ω
*fan and ω
*pump respectively, are provided as
inputs to VFD microcontrollers, which configure each inverter reference voltage, v
*ref,i. Sensorless vector control based on
voltage (vm) and current (im) measurements on motor side is used to drive the inverter semiconductor switches.
The applied VFD topology for each motor consists of a six-pulse diode rectifier and a standard two-level three-phase voltage source inverter (VSI) implemented with IGBTs. A 3% dc choke is included in the dc-link of all drives and an extra 3% ac choke is used for cooling sea water pump case. The required choke inductances, L, are calculated as shown in Eq. 14 with reference to the current flowing in the AC supply, ii,a, and to line-to-line supply voltage, vLL at fundamental frequency, f , to achieve a desired voltage drop across the reactor, ΔvL,ch, in order to satisfy relevant marine harmonic standards [21]:
ai
LLchL
ai
chL
If
V
If
VHL
,
,
,
,
~2100
)3/~(%
~2
~
][
(14)
B. Data-driven process control strategy for flow regulation
The speed regulation of E/R fan and C.S.W pump motors is performed according to a data-driven process control stategy to match the variable flow requirements of E/R ventilation and central cooling water systems according to M/E load. To this end, Table II shows how sea-going operation in reduced M/E load (60% of MCR) results in considerable variation of required air flow from fans, qfan, and S.W flow from C.S.W pump, qpump, for the vessel under study. It is shown that qfan,60%
is equal to 107,000m3/h, nearly 50% lower than designed air
flow due to less combustion and heat evacuation requirements for M/E and auxiliary machinery. Similarly, qpump,60% is equal to 258.4m
3/h, nearly 40% lower than the designed SW flow at
100% of M/E MCR. This occurs because the heat dissipation needs of E/R consumers in the central cooling water system associated with M/E and auxiliary machinery operation demand less capacity from the central cooler, i.e Qcent,60% is reduced by 42% compared to nominal capacity, Qcent,MCR. To exploit this power saving through proper VFD operation, the outline of the applied data-driven process control strategy for effective speed regulation is presented in Fig.4.
For speed regulation of E/R fan motors, two independent proportional-integral-derivative (PID) controllers are used to define the airflow required to maintain certain differential E/R overpressure (15-45Pa) and E/R ambient temperature (~36
oC),
resulting in respective q*fan,P and q
*fan,T air flow references.
Then depending on number of operating fans, Nfan, the final air flow reference for each fan, q
*fan,i, is defined as the maximum
between q*fan,P and q
*fan,T to satisfy E/R needs as per Eq.15:
fanTfansPfansifan Nqqq /},max{ *
,
*
,
*
, (15)
The corresponding speed reference for each fan motor and the resulting power consumption in comparison to nominal capacities are then determined by affinity laws as in Eq.16:
3
,
,
,
,
,
,
MCRfan
ifan
MCRfan
ifan
MCRfan
ifan
P
P
q
q
(16)
It is shown that the theoretical 50% reduction in air flow or speed reference experienced during sea-going operation with
Fig. 4. Outline of the proposed data-driven process control for flow
regulation in E/R ventilation and central cooling water system processes
M/E working at 60% of PMCR will result in about 85% reduction in fan power consumption. However, the final fan speed reference, ω
*ref,fan,i provided to the VFD is bounded by a
minimum speed, ω*min,fans according to Eq.17:
%95*
,,min, ifanreffans (17)
This limit is defined by the following process constraints: (i) keep scavenge air pressure higher than 0.7bar (ii) keep temperature in scavenge air chamber less than 48
oC and (iii)
keep motor speed is all cases higher than 40% of ωnom to avoid overheating. After the above checks which determine the fan VFD operation, their resulting power consumption should be considerably lower, leading to reduced D/G load.
In a similar fashion, for speed regulation of C.S.W pump motor two independent pairs of PID controllers are used to define water flow required during sea-going and discharging conditions respectively. At sea-going mode, where M/E is working, the controller maintains certain S.W pump discharge pressure on one side of central cooler (1.2-1.6bar) and F.W temperature on the other side of central cooler (34-36
oC),
resulting in respective q*pump,Psw and q
*pump,Tfw water flow
references. However, in cargo discharging condition, the vacuum condenser is in use together with the steam – driven cargo pumps, so the controller maintains certain S.W differential temperature between outlet and inlet of vacuum condenser (3.4-6.0
oC) together with constant S.W pressure,
resulting in q*pump,Tsw water flow reference. Therefore,
depending on number of operating C.S.W pumps, Npump, and vessel’s operating condition, the final pump water flow reference for each C.S.W pump, q
*pump,i, is always the
maximum between the two calculated references, as per Eq.18:
pumpPswpumpTswpumpocfan
pumpTfwpumpPswpumpseapump
ipumpNqqq
Nqqqq
/},max{
/},max{*
,
*
,
*
arg,
*
,
*
,
*
,*
,
(18)
The corresponding speed reference for each centrifugal pump motor and its resulting power consumption compared to the nominal are then defined by affinity laws as in Eq.19:
3
,
,
,
,
,
,
MCRpump
ipump
MCRpump
ipump
MCRpump
ipump
P
P
q
q
(19)
It is shown that the theoretical 40% reduction in water flow or speed reference experienced during sea-going operation with M/E working at 60% of PMCR will result in about 75% reduction in pump power consumption. However, the final pump speed reference, ω
*ref,fan,i provided to the VFD is
bounded by a minimum speed limit, ω*min,pumps as in Eq. 20:
%95*
,,min, ipumprefpumps (20)
This limit is defined by the following process constraints: (i) maintain M/E jacket cooling F.W outlet temperature less than 85
oC (ii) keep M/E jacket cooling F.W pressure more than
3.6bar (iii) keep D/G jacket cooling F.W pressure more than 3.4bar (iv) maintain F.W system temperature below 38oC and (v) keep S.W pressure in the outlet side of atmospheric condenser more than 0.5bar. After the above checks which determine pump VFD operation, the resulting power consumption should be considerably lower, leading to reduced D/G load, especially at sea-going.
IV. EXPERIMENTAL VERIFICATION AND DISCUSSION
A. Power consumption optimization for VFD-equipped loads
The proposed topology is applied as a retrofit installation to all four 18.5kW, 440V, 60Hz E/R ventilation fans and to one 75kW, 440V, 60Hz C.S.W pump in the marine tanker vessel under study. To evaluate the proposed VFD topology performance towards optimizing power consumption of E/R fan and C.S.W pump motors, the system response is initially tested during a typical 10-day voyage that includes loading, sea-going and discharging operations, as presented in Fig.5.
The power consumption variation of NO.1 E/R fan motor operating in variable frequency mode through the proposed topology is compared to its original nominal DOL power consumption of 18.5kW, as presented in Fig. 5(a). The power consumption waveform is the same for all the working VFD-equipped E/R fans. It is shown that while M/E is working at high load, four fans are configured to work at 43Hz to compensate for air flow requirements, resulting in 7kW power consumption for each fan, thus leading to a 46kW D/G power saving margin compared to previous four fan direct-on-line (DOL) operation. However, during sea-going operation with M/E working around 60% of PMCR at reduced speed, the controller configures either 4 fans to work at 30Hz or 3 fans to work at 43Hz. In that case, the respective individual fan power consumption is either 2.5kW or 7kW, leading to a considerable
(a)
(b)
Fig. 5. Evaluation of created D/G power saving margin from E/R fan and C.S.W pump motor operation after the proposed VFD topology application compared to original condition for a typical voyage: (a) Power consumption of E/R fan motor (NO.1 of 4) (b) Power consumption of C.S.W pump motor
D/G power saving margin in the range of 53kW to 64kW. It has to be noted that if less fans are working, the controller raises the speed of remaining online units to achieve required air flow references. In relation to the above, the power consumption variation of S.W cooling pump motor operating in variable frequency mode through the proposed topology is compared to its original nominal power consumption of 75kW, as presented in Fig. 5(b). It is shown that during loading period, where M/E is not working and E/R heat dissipation needs are minimum, the controller configures the pump to work at 42Hz (70% of speed), leading to pump power consumption of 23kW or 52kW power saving margin from D/G load. During sea-going mode, it is shown that while M/E operates in higher speeds above 70-75% of PMCR, the S.W pump motor speed is raised to 56-58Hz in order to maintain central cooling system dissipation needs. In this area, the S.W pump power consumption is 60-70kW, thus the power saving margin is small. Nevertheless, the effectiveness of the proposed topology is highlighted during slow-steaming sea-going operation, where M/E works at 55-60% of MCR and SW temperature is below 25
oC. In this area, it is shown that
the pump works at 42Hz (70% of ωnom), consuming 25kW instead of the nominal 75kW. The created power saving margin of 50kW is significant and positively affects the power balance during this period. Finally, system’s performance is also assessed during cargo discharging period, where Fig. 5(b)
shows that the pump is configured to work at 54Hz (94% of ωnom), thus consuming 60kW to compensate for the demands of additional E/R machinery such as the vacuum condenser and the cargo pumps.
B. Power balance and D/G fuel consumption optimization
Based on power saving margin resulting from proposed data-driven control of VFD-equipped E/R ventilation fans and C.S.W pump, the impact on power balance and D/G fuel consumption is evaluated hereby. To this end, the variation of total power system load and D/G fuel consumption is presented in Fig.6 before and after proposed VFD topology installation for a period of 45days, including anchorage, loading, sea-going (high speed and slow-steaming) and cargo discharging conditions according to data obtained from the vessel.
During sea-going period before VFD application, it is shown that while the vessel was travelling at 11.0knots and its M/E was operating around 55% to 60% of PMCR, the average power system load to be serviced by D/Gs was between 650-670kW or around 82-85% of Pr,gi, exceeding the PMS load limit. Thus a second D/G had to be used in parallel and perform load sharing at 330kW. This fact would lead to D/G SFOCMDO being equal to 208.17g /kWh, as derived from Fig.1 or to an equivalent SFOCHFO of 225.03g/kWh since Heavy
Fig. 6. Variation of total power system load and D/G fuel consumption for a
period of about 45days (21/04/2013 to 10/06/2013) before and after the
proposed VFD topology installation for typical vessel operation (a) Vessel’s M/E %MCR, speed (knots) and operating condition during this period (b)
Comparison of vessel’s total power system load before and after the proposed
system implementation (c) Variation of daily D/G fuel consumption compared to the previous operating practice
Fuel Oil (HFO) is used with lower calorific value (LCV) than MDO. As shown in Fig.6(c), this leads to daily D/G fuel consumption (FOC) of about 3.1tons. The slow-steaming operating practice at sea-going occupies around 46% of the vessel’s annual trading time, so the impact of proposed system to D/GFOC is expected to be substantial during this period.
This expectation is verified by ship’s sea-going operation results obtained after the proposed system retrofit installation. Fig. 5(b) shows that at the same M/E load as before, the average power system daily load has been reduced to about
Fig. 7. Systematic procedure for calculation of economic benefits in relation
to vessel’s operating profile, D/G load and fuel consumption
535kW or 68.6% of Pr,gi, due to power saving compensation created by the topology used to drive E/R fan and C.S.W pump motors, which results in total load reduction of about 125kW or about 19%. This means that now only one D/G is sufficient to service the required load at slow-steaming operation, which brings significant benefits. First and foremost, the D/G SFOCMDO is now reduced to 193.44g/kWh, or equivalently to SFOCHFO equal to 209.11g/kWh, since one D/G is now working at much higher load. This results to daily average D/G FOC of about 2.5tons, therefore a gain in daily fuel consumption of about 0.5tons. Secondly, the maintenance costs are greatly reduced since the D/G running hours are also reduced. A second D/G is needed only in case some intermittent loads are in service, such as operation of ballast pumps, greatly improving vessel’s power balance.
While this power saving margin greatly affects the vessel’s power balance during sea-going operation at low M/E load, it is not the same for other vessel operating conditions. It grows during anchorage/ loading because M/E is not working and heat dissipation needs of auxiliary machinery are less and shrinks during sea-going operation at high load and during discharge operation. As shown in Fig.6, the expected total power system load during discharge operation without the proposed topology was about 1014kW, attributed to two D/Gs working in 65% load each. After the proposed topology application, the load during discharging has been reduced to 940kW, still needed to be serviced by two D/Gs working at 60% each. Taking into account that low sulfur MDO is usually used at discharging ports, the SFOCMDO remains around 195g/kWh before and after the application of the proposed
Fig. 8. Comparison of vessel’s annual D/G fuel consumption throughout its
operating profile before and after the proposed topology retrofit application
Fig. 9. Comparison of vessel’s annual energy consumption throughout its
operating profile before and after the proposed topology retrofit application
topology, thus leading to fuel savings only from absolute load reduction. The load variation is performed in similar fashion during sea-going with M/E working at high speed as shown in Fig.6. Nevertheless, sea-going operation at high M/E load and cargo discharging are only 10% of the vessel’s trading time, so the effects of the proposed topology are limited.
C. Techno-economic assessment method
To define the relation of power balance and D/G fuel
consumption optimization to the obtained economic benefits
according to vessel’s operating profile and type of fuel used, a
systematic procedure is proposed in Fig. 7 and applied in this
work. Based on vessel’s operating condition, there is a specific
power system load demand from D/Gs, PL,g. This load demand
is optimized by the proposed topology. Based on PL,g, and the
constraints described in the first section the PMS assigns a
number of D/Gs which perform load sharing if needed, with
PL,gi load assigned to each D/G. Then based on (10), the D/G
mechanical power, Pmech,gi, is calculated and used to obtain the
corresponding SFOCMDO value from D/G SFOC curve for
MDO use. The next step is to proportionally correct SFOC for
HFO use if needed. In that case SFOCHFO should be
compensated for HFO LCV as in Eq.21:
TABLE III
SUMMARY OF ANNUAL BENEFITS FOR THE VESSEL UNDER STUDY
500,39
700,42MDO
HFO
MDOMDOHFO SFOC
LCV
LCVSFOCSFOC (21)
Based on SFOC, D/G FOC can be calculated in tons according to (9). The corresponding GHG emissions can then be calculated according to (11)-(13). The economic benefits can be calculated from FOC based on fuel prices.
D. Economic benefits from the proposed topology
Based on the experimental results, on data acquired from the vessel and on the techno-economic assessment method mentioned in the previous section, the economic benefits of vessel’s power system energy and D/G fuel consumption are
calculated throughout the vessel’s operating profile. Results comparison is made before and after the proposed topology retrofit installation. More specifically, Fig. 8 shows the annual D/G fuel consumption and Fig. 9 presents the annual energy consumption and the respective economic benefits. It is taken for granted that always HFO is used throughout sea-going operation, 70% HFO and 30% MDO are used during port stay/loading operation, and MDO is always used during cargo discharging operation. 2013 and 2014 fuel prices are considered for all economic benefit calculations.
From Fig. 8 it is evident that the biggest benefit occurs during sea-going operation in slow-steaming condition, where M/E operates in the range between 55-60% of MCR. Results show that the annual savings in energy and D/G fuel consumption at sea-going condition are 557.3MWh and 163tn of HFO, validating the impact of the proposed system. Table III summarizes the annual benefits for the vessel under investigation.
V. CONCLUSIONS
In this paper, a topology was developed to optimize D/G operating profile and fuel consumption in autonomous power
systems of marine vessels, especially during sea-going period with M/E working at lower speed. It was shown that load variation compared to sea-trials has affected onboard power balance, often requiring PMS to run an additional D/G during sea-going operation to satisfy operational constraints. To this
end, a process control topology involving VFD application to E/R fans and C.S.W pump was proposed to exploit the power
saving margin created at slow-steaming due to reduced air and cooling water flow requirements of M/E and auxiliary machinery. The proposed topology was implemented as a retrofit installation in a typical tanker vessel. It was theoretically and experimentally shown that by compensating the load variation at sea-going with the created power saving margin from the proposed topology, the constant use of only one D/G is achieved, significantly optimizing power balance, fuel consumption and GHG emissions. It was shown that total annual energy savings approach 1000MWh, while annual D/G fuel consumption is reduced by 183tons HFO and 52 tons MDO. Additionally, annual emissions of CO2, SO2 and NOx
are reduced by 748tons, 7.5tons and 18.6tons respectively.
ACKNOWLEDGMENT
This work was supported by Thenamaris Ships Management Inc under its strategic program for energy management optimization across its fleet.
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During recent years, Silicon Carbide (SiC) power semiconductor device technology has become increasingly attractive for use in power converters over Si due to features that include operation in high temperatures, high voltage blocking capability, thermal conductivity, switching speed, low on-state resistance and low voltage drop [1]-[6]. In particular, JFETs are known to be highly mature among SiC semiconductor devices with normally-on depletion mode (DM) variants possessing the most favorable characteristics for use in high efficiency converters [7]-[9].
Whereas robust gate drive solutions have been presented for normally-off SiC JFET equivalents [10]-[11], the normally-on feature of DM SiC JFETs leads to critical challenges when these power devices operate in a power converter bridge leg configuration. Due to their normally-on feature, a negative
gate-to-source voltage, vgs, lower than pinch-off voltage, Vp, is required to turn-off the device. Therefore, a gate drive malfunction, a mismatch in gating signals or coupled noise through the Miller capacitance, Cgd, could lead to cross
conduction in an inverter bridge leg, causing shoot-through fault. To address the above, the gate drives proposed in [12]-[13] use an RCD network to operate the device in zero bias during on-state and close to its breakdown voltage during off-state. In these approaches, however, the lack of forward bias results in increased conduction losses, especially if pulse width modulation (PWM) techniques are used, since applied vgs may be negative during on-state while pulse width decreases. Moreover, the converter reliability is still vulnerable to overcurrent or shoot-through faults in one of its phase legs, caused by gate drive component or power supply failure.
To deal with these constraints, overcurrent protection methods based on desaturation technique are described and evaluated in [14]-[15]. The drawbacks of these solutions are that shoot-through fault can still occur in case of logic unit, driving circuit or power supply malfunction, while fault response time exceeds 10μs. To this end, a technique dealing with total gate drive power supply failure is presented in [16], where an IGBT and a relay are combined to form a bidirectional switch embedded in the converter dc-link. However, depending on input voltage, inductances and breaker response time, the SiC JFETs are still subject to overcurrent condition from the dc-link loop. Another protection approach to secure inverter operation at power-up without gate drive power supply is proposed in [17], where the impact of JFETs’ aging in the VSI is also evaluated. While the safety system is capable of running in elevated temperatures as well, it is configured to be applied only to SiC inverters. The concept of self-powered gate drive not depending on power supply failure, is introduced in [18]-[19], requiring the design of an LC resonant circuit in a start-up converter, operating for the first quarter of the first time
period. It is shown that a trade-off between performance and stability is required.
In this work, a standalone ac-coupled gate drive with integrated protection functions against overcurrent and shoot-through faults is developed and analyzed for use in power converters built with normally-on SiC JFETs. The gate
(a)
(b)
Fig. 1. Overview of the proposed self-protection topology integrated into the
gate drive for normally-on SiC JFETs (a) Layout of the proposed topology (b)
Laboratory prototype used for experimental verification
drive applies forward bias during on-state, achieving significant conduction and switching loss reduction in comparison to zero bias gate drives. Using a variation of the common RCD network in the output stage, gate-to-drain interactions that induce coupled noise are minimized. Apart from faults associated with distortion of vgs waveform, the proposed gate drive topology deals with power supply failure and overcurrent conditions by turning the device off within its specified fault withstand time (around 50μs). More specifically, the desaturation method is used to monitor drain-to-source voltage, vds, in order to detect overcurrent conditions based on the device saturation characteristics. An overcurrent protection topology clears the fault in less than 5μs by applying negative vgs, thus forcing transition to the off-state. In case of gate drive power supply failure, an additional configuration is activated to keep the device in off-state for sufficient time until converter isolation.
The integrated gate drive and protection topology, which is presented in Fig. 1(a) and implemented in a laboratory prototype, as shown in Fig 1(b), was built with the intention of minimizing parasitic elements. To this end, the topology was built with one layer dedicated to ground plane in order to minimize stray inductances. Throughout the study, the device
used is a TO-247 package, 1200V, 27A normally-on SiC JFET, model SJDP120R085. A double pulse tester set-up is used to experimentally verify the performance of the proposed protection schemes under various fault conditions as well as measure conduction and switching losses when the device operates in a converter bridge leg configuration.
II. PROPOSED GATE DRIVE CIRCUIT USING FORWARD BIAS
The device driving section of the proposed gate drive circuit is shown in Fig.2. The isolation stage, implemented with an optical isolator of reduced propagation delay, provides ohmic isolation between the converter control circuit and the gate drive. The embedded power supply and voltage regulator common nodes are connected to the respective JFET source, which is usually subject to fast changes in potential, enhancing the common mode immunity of the system and eliminating the need for using common mode chokes. The power stage consists of a totem pole N-MOS and P-MOS configuration that applies forward bias during conduction, +VCC, or negative bias at turn-off, -VSS, to the device gate-to-source terminals. The parameters of the output stage, consisting of three parallel R-C-D branches, are designed to optimize performance and minimize coupled noise during switching. The decoupling capacitor, Cg, provides a path to ground for parasitic current icgd during switching, by improving the current divider in the JFET gate terminal. Design considerations for each parameter are made by investigating the operating phases of the proposed driving circuit presented in Fig.3.
A. Device on-state operation
To reduce on-state resistance, Rds-on, a positive bias, +Vcc, is applied across gate-to-source terminals in order to keep the device turned-on. As shown in Fig. 3(a), the gate current, ig, flows through the parallel branch consisting of the Schottky diode Df and the series resistance Rf. According to the device ig - vgs characteristics at forward bias, Rf is selected according to the desired operating point [Vgs-fwd, Ig-fwd] during on-state condition, so that forward gate current ig-fwd is limited to a
desired value with reference to the voltage drop across Df, VDf, as expressed in Eq. 1:
fwdg
DffwdgsCC
fI
VVVR
,
(1)
To minimize gate drive power losses, the desired Ig-fwd is set at 20mA during on-state, corresponding to Vgs-fwd equal to 2.67V. Based on (1), Rf is set at 100Ω. At the same time, capacitance Cac of the upper branch is charged due to voltage difference vCac developed across its terminal during on-state, as shown in Eq. 2:
fwdgsCCCac V (2)
In this operating phase, the conduction energy losses depend on the conducting inverval, ton-state, as expressed in Eq.3:
stateonondsdstateonddsstateon tRitiW 2 (3)
Fig. 2. Proposed ac-coupled gate drive circuit for normally-on SiC JFETs, applying forward bias during device conduction
Fig. 3. Operating phases of the proposed gate drive circuit (a) Device on-
state operation (b) Device on-to-off state transition (c) Device off-state
operation (d) Device off-to-on state transition
B. Device on-to-off state transition
As shown in Fig. 3(b), during device turn-off transition, a
negative voltage, -VSS, is applied to gate-to-source terminals through the upper branch, consisting of Cac and Rac. Nevertheless, since Cac was previously charged, the applied vgs is more negative than –VSS as expressed as in Eq. 4:
CacSSgs V (4)
The series resistance, Rac is selected so that the discharging time of the device intrinsic capacitance Cgs is as small as possible in the range 3-10Ω, while at the same time vds and id oscillations are minimized. For proper value selection, the power stage output resistance and device gate resistance, RG, has to be considered. In this application, Rac is chosen as 5Ω.
Moreover, the capacitance Cac is calculated as in Eq. 5:
Cac
CgCac
acV
QQC
(5)
where QCac is the charge equal to the one required by the device during turn-off switching, QSiC, while QCg is the charge
of the external capacitance Cg, used for minimizing coupled noise. Based on (2) and (5), capacitor Cac should be selected within the range defined by Eq. 6:
fwdgCC
SiCg
ac
fwdgCC
SiCgSS
VV
QQC
VV
QCV
2 (6)
Since the operating point is chosen at the knee of Ig-Vgs characteristic as shown in Fig.4, the limits obtained for Cac are 41.6nF and 65nF according to (6). Therefore, Cac is chosen to be 47nF to achieve optimum performance. Switching losses during on-to-off transition depend on turn-off time, toff, and are calculated by Eq.7:
offt
ddsoffsw dttitW0
)()( (7)
C. Device off-state operation
To limit potential gate breakdown current, Ibr, which may flow in case vgs greatly exceeds Vp with temperature increase, the circuit consisting of a fast recovery Schottky diode, Dp, and a high value resistor in series, Rp, is included in the output stage, as shown in Fig.3(c). The value of Rp is defined by Eq.8
in order to limit the potential avalanche current:
br
gsSS
pI
VR
(8)
A 3kΩ resistor was chosen for this application. Provided that
the normally-on SiC JFET does not enter breakdown region,
the decrease of Rp contributes to dealing with gate-to-drain
interactions, since a low resistance path is formed for the
parasitic current icgd, which flows towards –VSS.
D. Device off-to-on state transition
As shown in Fig. 3(d) during turn-on transition, forward bias, +VCC, is applied to the gate through the output stage branch consisting of Rac and Cac, which is fully discharged from
(a)
(b)
Fig. 4. Normally-on SiC JFET device under test (DUT) (a) ig-vgs
characteristics at forward bias with selected operating point during conduction (b) id-vds saturation characteristics at various temperatures (25oC, 100oC, and
150oC)
the previous operating condition. The switching losses during off-to-on transition can be calculated by Eq. 9:
ont
ddsonsw dttitW0
)()( (9)
The gate driver component selection for the laboratory prototype used in the experiments is presented in Table I.
III. PROPOSED GATE-DRIVER SELF-PROTECTION TOPOLOGY
To protect against overcurrent condition, gate drives built for conventional bipolar semiconductor devices (e.g IGBTs) monitor the current rise and trigger a protection circuit when
TABLE I
COMPONENT SELECTION FOR GATE DRIVE CIRCUIT (FIG.2)
the device pulls out of saturation, resulting in increased collector-emitter voltage. However, unlike most Si devices featuring high impedance constant current active region, transition from ohmic to active region for SiC JFETs is not clearly defined, requiring a different approach. Furthermore, under higher switching speeds, the noise immunity of desaturation circuit should be enhanced in order to avoid false triggering of the protection circuit, while maintaining a fast response time of about 5μs.
The desaturation protection method implemented hereby involves monitoring of voltage drop across drain-to-source terminals, vds, only during conduction of SiC JFET. During normal operation, vds remains less than few hundred mV. When an overcurrent condition occurs, vds increases rapidly following a temperature dependent saturation curve, triggering an overcurrent protection circuit integrated into the gate drive. Regarding the normally-on SiC JFET under study, the ig-vgs characteristics are presented in Fig. 4(a), showing the selected forward bias operating point during conduction phase [Vgs-fwd,Ig-fwd]=[2.67V, 20mA]. In Fig. 4(b), the id - vds
characteristics are provided for temperature range of 25oC-
150oC. It is shown that for a given drain-to-source voltage
drop, the corresponding drain current fault threshold, id, highly depends on ambient temperature. In particular, for a 2V vds
voltage drop level during fault, the corresponding drain current would be 32A at 25
oC, 18A at 100
oC and 14A at
150oC, indicating the need to define vds level that triggers the
protection circuit in relation to ambient temperature.
A. Design of the gate-driver self-protection topology
The design layout of the proposed self-protection topology interacting with the gate drive circuit for normally-on SiC JFETs is presented in Fig. 5. More specifically, it includes two
circuits with protective functions against overcurrent or shoot-through fault conditions during normal gate drive power
supply operation, as shown in Fig. 5(a), and protective functions against gate drive power supply failure, as presented in Fig. 5(b).
1) Overcurrent protection based on desaturation method: The overcurrent protection circuit based on desaturation fault detection method that is implemented in this work is presented in Fig. 5(a). A high voltage fast recovery diode, Dss, whose anode is positively biased with +Vcc, is connected to drain terminal of the normally-on SiC JFET and thus conducts only when the semiconductor switch is in on-state. Once anode
(a)
(b)
Fig. 5. Design of the proposed self-protection topology interacting with the
gate drive for normally-on SiC JFETs (a) Proposed overcurrent protection
circuit based on desaturation fault detection method (b) Proposed gate drive
power supply failure protection circuit
voltage is filtered from spikes and parasitic noise through the low pass filter comprising of Rin and Cin, the signal vdesat is compared to reference voltage, Vref, using a comparator which features fast response time. The fault response time relies on blanking time delay required for vds to reach its fault triggering value following the specific temperature-dependent saturation curve, while vdesat is a function of vds, as in Eq. 10:
ssonDssindsdesat iRRtt )()()( , (10)
where RDss,on is the Dss on-resistance and iss is the current flowing through this diode during conduction state.
The reference voltage, Vref, is set by resistors Rref1 and Rref2 to define the level of desired overcurrent drain current limit, Id,limit, where the comparator would produce a fault signal. To ensure that comparison of vdesat and Vref signals takes place only while the SiC JFET is in the on-state, the circuit that
includes N-Mosfets Mcomp and Mgd interfaces with the gate drive (DRIVE signal) to impose –Vss to the non-inverting input of the comparator during off-state, avoiding any false fault triggering at this period. More specifically, when the gate drive outputs negative voltage –Vss, lower than Vp, to turn-off the SiC JFET, Mgd is cut-off and Mcomp turns-on, leading to Vdesat equal to –Vss. On the other hand, while SiC JFET is in the on-state, the gate drive outputs +Vcc, turning-on Mgd and turning-off Mcomp, so that vdesat follows JFET saturation curve.
Nevertheless, effective synchronization of vgs with vdesat signal during transients is required. In particular, if there is delay in vdesat transition towards low potential, -Vss, and vgs has already reached –Vss due to SiC JFET blanking time, a false error signal will be triggered by the comparator. To overcome this issue, a filter formed by Rdel, Ddel and Cdel was designed to introduce the desirable delay in the turn-off transition of Mcomp, while simultaneously accelerate its turn-on process according to time constants described in Eq.11:
deldelgdoffMcomp
delgdonMcomp
CRRt
CRt
)( 2,
1, (11)
When a fault error signal is triggered by the comparator, it is received by the logic control unit, which includes an RS latch circuit. Its role is to flip and hold at high level output when a fault is detected. This signal is normalized to 3.3V through a zener diode, Zsig, to be provided as a flag signal to the SiC JFET gate signal controller in order to command removal of gating signals. The output of the RS latch is also used to drive the overcurrent protection circuit at the next stage, the role of which is to impose negative potential equal to –Vss to SiC JFET gate-to-source terminals in order to effectively turn-off the device in case an overcurrent fault is detected. To this end, an optocoupler acts as a buffer between the logic unit and the overcurrent protection circuit, while at the same time outputs 0V in case a fault is detected (its output is –Vss if no fault appears). A zener diode with suitable breakdown voltage, Zpr, whose anode is negatively biased at –Vss, regulates negative bias at node C at such level, so that the N-Mosfet, Mpr, is turned-on in case of fault, enforcing negative voltage –Vss at SiC JFET gate-to-source terminals from the gate drive power supply through a fast recovery diode, Dpr. A reset signal acting on the logic unit RS latch must be provided by the user after the overcurrent fault is detected in order to restore proper system operation. It is vital that the overcurrent protection scheme is designed so that fault detection and clearance is made in less than 5μs to limit the rise of fault current to acceptable levels.
2) Gate drive power supply failure protection circuit: The
proposed overcurrent protection scheme based on the
desaturation method functions properly only when gate drive
power supply remains constant and adequate. In case of power
supply failure, the gate driver will not be able to turn-off the
normally-on SiC JFET and the overcurrent protection scheme
will not be able to impose negative voltage to gate-to-source
terminals to keep the device in the off-state, causing shoot-
through fault in converters including phase leg configurations,
such as voltage source inverters (VSIs).
TABLE II
COMPONENT SELECTION FOR OVERCURRENT PROTECTION CIRCUIT-FIG.5(A)
To address this issue, a power supply failure protection circuit, presented in Fig. 5(b), is integrated into the gate drive and keeps the device in the off-state for sufficient time for the converter to be isolated from its power source. The level of gate drive power supply, +Vss, is monitored by a Zener diode, Zsup, whose breakdown voltage is selected to establish the desired level of acceptable power supply voltage. A combination of optocouplers is used to control the conduction of a N-Mosfet switch, Mdis, which in turn controls the discharge of an electrolytic capacitor, Cf, to the gate-to-source terminals of the normally-on SiC JFET, which remains in the off-state until the converter is isolated.
During normal gate drive power supply operation, Mdis is in the cut-off region, while the zener diode, ZD, whose anode is negatively biased with –Vss and its cathode is connected to ground, regulates the gate-to-source voltage of N-Mosfet, Mch, to be about 5V. Since the source of Mch is connected to –Vss and capacitor Cf is not charged, it is valid that Vds,Mch < Vgs,Mch = 5V, so Mch initially conducts in the ohmic (or triode) region. Since the diode, Dch, is conducting as long as Mch is conducting (Vgs,Mch > 0V), capacitor Cf is charging with voltage expressed as in Eq. 12:
tCRRR
fM,chchd,ch
SSCfchMchchd
fe
CRRRVt
)(
1
,,
)(
1)( (12)
where Rd,ch and RM,ch are the on-state resistances of Dch and Mch
respectively. Rch defines the charging speed of Cf ; the higher
is Rch, the slower is Cf charging speed.
In case of inadequate gate drive power supply, the zener diode, Zsup, regulates the voltage level at the input of the first optocoupler so that its output is –VSS. The gate voltage of N-Mosfet, Mdis, is then vg,Mdis = 0V, while the negative terminal of the previously charged electrolytic capacitor, Cf, is connected to the source of Mdis, leading to Vgs,Mdis = +VSS, which turns it on. Therefore, Cf discharges at the SiC JFET gate-to-source terminals through the path formed by Mdis, and
TABLE III
COMPONENT SELECTION FOR GATE DRIVE POWER SUPPLY FAILURE
PROTECTION CIRCUIT-FIG.5(B)
Ddis, and its voltage during this period is expressed by Eq.13:
0),1()0()()(
1
,,
tet
tCRR
CCfdisddisM
f (13)
where vC(0+) is the initial capacitor charge when gate drive
power supply failure occurs, while RM,dis and Rd,dis are the on-
state resistances of Mdis and Ddis respectively. At the same
time, a signal called EN is sent to the enable input of the gate
drive power stage IC to block the gating signals and another
signal is sent to the converter control circuit which provides
the command for the converter to be isolated. The discharge of
Cf via Dch is prevented during power supply failure, since Mch
is in the cut-off region.
B. Implementation of proposed self-protection topology
For the laboratory prototype used in the experiments, component selection for the overcurrent protection circuit presented in Fig. 5(a) is made as shown in Table II. To achieve minimum delays for fault detection and clearance, the propagation delays for the major components, including the sensing diode, comparator, logic unit and optocoupler were chosen 120ns, 200ns, 100ns and 300ns respectively. Similarly, component selection for the gate drive power supply failure protection circuit implemented in the prototype is made as shown in Table III.
IV. EXPERIMENTAL RESULTS AND DISCUSSION
A. Experimental set-up and testing conditions
The performance of the proposed gate drive topology is investigated using the experimental set-up presented in Fig. 6(a). It is a double pulse tester, comprising a 600Vdc source, a 350μH inductor load and two freewheeling SiC diodes. Two identical gate drive topologies are used to supply the upper normally-on SiC JFET and the DUT, interfacing with the control unit that provides the gating signals to the gate drive prototypes. The control unit, implemented with a TI F28335 digital signal processor (DSP), receives the input signals for power supply failure and overcurrent conditions in order to remove the gating signals and isolate the inverter.
(a)
(b)
Fig. 6. Experimental set-up and testing conditions of proposed gate driver
circuit (a) Double pulse tester (b) Peak currents applied to the SiC JFET gate
during turn-on and turn-off
To achieve fast gate drive response times under high switching frequencies, high current peaks were applied at the gate of the normally-on SiC JFET during turn-on and turn-off transients, which in turn affect the response of vgs. As shown in Fig.6(b), for a switching frequency of 2MHz, the gate current peaks, ig,p-p, applied by the gate drive during device turn-on and turn-off transients are 1.1A and 1.5A respectively with a pulse width of 60ns. The corresponding switching times for vgs are ton = 35ns during turn-on and toff = 60ns during turn-off. As shown in Fig.6(b), a positive bias of Vgs-fwd = 2.67V is applied during on-state according to the operating point selected in Fig. 4(a), leading to 7% reduction in conduction losses. Moreover, while turn-off voltage provided by the gate drive is -15V, vgs reaches -18V according to (4), before Cac is fully discharged.
B. Gate drive performance evaluation during transients
The gate drive response during device turn-on process is evaluated from the SiC JFET turn-on characteristics presented in Fig.7(a) for Vdc = 600V, switching frequency, fsw, equal to 25kHz, and 10A load current. It is shown that di/dt for drain current, id, is equal to 0.5kA/μs, resulting in a very fast turn-on time, ton, equal to 20ns. Similarly, vds rises from 0V to 300V with dv/dt equal to 15kV/μs. It is also shown that the design of gate drive results in vgs waveform being free of oscillations that
(a)
(b)
Fig. 7. Performance evaluation of proposed gate drive during transients (fsw =
25kHz, Vdc = 600V, Id =10A) (a) Switching performance evaluation at turn-on
(b) Switching performance evaluation at turn-off
would induce potential gate punch-through or shoot-through faults. According to (9), the switching losses at turn-on depend on turn-on time, ton, where crossover between vds and id waveforms occurs. Based on the experimental results, zero voltage switching (ZVS) takes place, since id begins to rise after vds has almost reached zero. This event leads to practically zero switching losses at turn-on, Wsw-on, which is also visible from switching losses waveform included in Fig. 7(a). The quality of vds and id waveforms during transients is also improved, featuring low overshoots and fast settling times.
The evaluation of gate drive response during the device turn-off process is carried out by examining the SiC JFET turn-off characteristics presented in Fig.7(b) for Vdc equal to 600V, fsw equal to 25kHz and 10A load current. It is shown that di/dt of id is equal to 0.5kA/μs, resulting in a very fast turn-off time, toff, equal to 22ns. Similarly, vds drops from 300V to 0V with dv/dt equal to 13.6kV/μs. It is also shown that vgs waveform is free of oscillations and reaches -18.5V before settling at -15V to sustain the device at off-state. According to (7), the switching losses at turn-off depend on turn-off time, toff, where crossover interval between vds and id
Fig. 8. Experimental evaluation of gate drive overcurrent protection circuit
response, using the implemented desaturation fault detection method, under a shoot-through fault condition (Vref = 3V, Id,limit = 45A, 25oC). Waveforms
include vds, vgs, id and comparator voltage signal, vdesat
waveforms occurs. Based on the experimentally obtained transients, the switching losses distribution waveform at turn-off is derived and presented in Fig.7(b). It is shown that switching losses at turn-off, Wsw-off, are limited to 180μJ, verifying the driving circuit performance.
C. Performance evaluation of overcurrent protection scheme
based on desaturation method
To validate the proposed overcurrent protection scheme performance, a shoot-through fault is caused in the phase leg implemented with SiC JFETs. This is performed by providing a single pulse that turns-on the upper device while the lower DUT is switching at 25kHz. The objective is to measure response time, tr, until the fault is cleared as well as the maximum value that the fault current, Id,max, reaches before it is limited and cleared by the protection circuit. For that purpose, the desired overcurrent limit of drain current is set at Id,limit = 45A, which will yield a drain- to-source voltage of Vds,limit = 3V, according to device saturation curve that
corresponds to the 25oC characteristic, as shown in Fig.4(b).
Therefore, the voltage level of signal vdesat provided to the comparator non-inverting input should rise from few mVs (during normal operation) up to a voltage level higher than the Vds,limit according to (10). Thus, in order to detect and clear the specified level of overcurrent, Id,limit, the comparator reference voltage, Vref, is set via Rref1 to 3V.
Experimental results presented in Fig.8 involve response to overcurrent condition for Vdc = 400V and SiC JFET operation at 25
oC. Before the fault occurs, it is shown that signal vdesat
swings between two values; from a few mV during JFET conduction to -15V while the SiC JFET is in the off-state. This verifies that fault detection takes place only during JFET on-state, while proper synchronization with vgs contributes towards avoid false protection circuit triggering. When the shoot-through fault occurs, id rises with very high di/dt, leading to vds rising according to the device saturation curve at 25
oC, which is sensed by diode Dss. At the exact moment
where specified Id,limit is reached, it is shown that vdesat has risen by ΔVdesat = 4V, triggering a fault signal which is provided to the converter control circuit in order to stop gating signals as well as to the overcurrent protection circuit that imposes -15V to the device gate-to-source terminals from the gate drive power supply. The protection topology response time, tr, is defined from the moment the drain current exceeds the specified limit of Id,limit = 45A until vds becomes equal to Vdc and the fault is cleared. Thus it is experimentally shown that overcurrent condition lasts for tr = 3.5μs, while the maximum level that fault current reaches within this period is Id,max = 70A. Both values are well within limits that the semiconductor device can withstand, verifying the effectiveness of the approach. It is also shown that while the fault is detected within 3.0μs, id drops down from 75A to zero within 0.5μs, validating the overcurrent circuit response. Afterwards, the overcurrent protection circuit keeps the device in the off-state until a reset signal is provided to the logic unit by the user to restore operation.
D. Performance evaluation of protection scheme under gate-
driver power supply failure condition
In order to extract the performance characteristics of the protection circuit presented in Fig.5(b), total power supply failure is caused by switching-off the gate drive power supply while the device operates at 25kHz with Vdc =100V. The response characteristics of the proposed protection circuit are presented in Fig.9. The objective is to evaluate how fast the switch is turned-off as well as the time frame allowed for the converter to be successfully isolated.
From the zoomed view of Fig.9, it is shown that when the gate drive power failure occurs, vp-fail signal measured at SiC JFET gate changes from zero to -15V within less than 0.2μs, since Mdis is turned-on and the previously charged capacitor, Cf, is discharging at the device gate through Ddis. The signal EN is at the same time triggering the enable function of the gate drive power stage IC which blocks the gating signals. From the zoomed out version of the response, it is shown that vgs follows the vp-fail signal according to Cf discharge curve, so the SiC JFET remains in the off-state for 16.8s, which is sufficient time for converter control circuit to command the isolation of power converter from its main power supply. The higher the capacity of the discharging capacitor, the longer is the available time frame for successful converter isolation. However, thorough monitoring of the capacitor condition is important for the reliability of the proposed circuit.
V. CONCLUSIONS
In this paper, a fast forward-biased gate drive circuit
featuring protection functions against overcurrent and gate
drive power supply failure conditions was designed and
implemented for normally-on SiC JFETs used in power
converter applications. Significant design considerations were
discussed for each proposed solution. A laboratory prototype
was built to evaluate the driving and protection schemes. From
the experimental results, it was shown that switching between
operating states takes place within 20ns, achieving transients with minimized overshoot and settling times, while conduction losses are also reduced. The total switching losses
Fig. 9. Experimental evaluation of gate drive power supply failure protection circuit response, following a power supply failure at 25kHz. The waveforms
presented are gate-to-source voltage vgs, capacitor discharge voltage at SiC JFET terminals, vp-fail, drain current, id, and drain-to-source voltage, vds
are limited to 180μJ, while ZVS occurs at turn-on. The
protection functions address all faults associated with the
normally-on characteristic. More specifically, the shoot-
through fault is detected using the desaturation fault detection
method and is cleared within 3.5μs, while the overcurrent
protection circuit can be easily parameterized to act at a
specific fault current level. In case of gate drive power supply
failure, a time frame of 16.8s is provided to effectively isolate
the power converter from its main power.
ACKNOWLEDGMENT
This work was financially supported by the European 7th
Framework Research Program “ANTI-SiC 09 SYN-32-1181”
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Page 1 of 11 2013-ESC-223
Development of an integrated energy efficiency control system for ship power balance and diesel generator fuel consumption optimization
Spyridon V. Giannoutsos Student Member, IEEE
Dep. of Electrical and Computer Engineering National Technical University of Athens (NTUA)
Abstract - In the maritime industry, high fuel costs cause
changes to the vessel's operating profile regarding the use of
Main Engine (M/E) and auxiliary machinery. This results in
deviations from the original load balance study and Power
Management System (PMS) adjustments that were performed
during shop tests, as Diesel Generators (DIGs) often operate in
parallel at low load far from their best efficiency points. In
this paper, an integrated process control scheme is proposed
in order to improve power balance in the vessel's air
ventilation and central water cooling systems, involving the
variable frequency control of Engine Room (E/R) Fans and
Cooling Sea Water Pumps (C.S.W PIPs). Experimental results
derived from system retrofit application to a typical 105,000 DWT Aframax tanker vessel verify the effectiveness of the
proposed control scheme towards achieving significant energy
savings, which are in accordance with the theoretical results
obtained from the air and heat balance studies. Power balance
improvement leads not only to fuel savings but also to the use
of only one DIG while sea going. Power quality measurements
show compliance with relevant marine harmonic standards.
Index Terms-Energy management, Marine vehicle power
systems, Diesel engines, Industrial power system control,
Industrial power system harmonics, Process control, AC-AC
power conversion, Variable speed drives, Pumps, Flow control
I. INTRODUCTION
A promising method in order to reduce fuel consumption in the maritime industry is to optimize the power balance by adapting to the vessel's actual power requirements [1]-[3]. However, the operating profile of Diesel Generators (DIGs) in the vessel's electrical distribution network is adjusted through startlstop tables during sea trials and is configured by the Power Management System (PMS), which determines the maximum limits in continuous and intermittent loading in order to prevent blackouts. Due to high fuel costs, vessels nowadays use slow steaming in order to operate Main Engine (M/E) at speeds lower than its specified Maximum Capacity Ratio (MCR) [4]. This practice, however, results in using DIGs with increased Specific Fuel Oil Consumption (SFOC) especially during sea going period, since changes in Engine Room (EIR) motor operation are not compliant with the original load balance study and PMS settings. As a result of slow steaming practice, the requirements of MIE and auxiliary machinery regarding combustion air and cooling water flow are much lower when compared to capacities required at maximum MCR.
Stefanos N. Manias Fellow, IEEE
Dep. of Electrical and Computer Engineering National Technical University of Athens (NTUA) 9, lroon Polytechniou st. 15780 Athens, Greece
Therefore, there is great opportunity to improve the vessel's energy efficiency by applying Variable Frequency Drives (VFDs) in order to adjust the speed of centrifugal fans and pumps related to the processes involved [5]-[6].
In the past, technical and economical aspects as well as the performance evaluation of VFD application to electric motors have been thoroughly reviewed in [7]-[9]. Different system design options for flow regulation in pumping systems and their economic impact have been discussed in [10]-[12], while in [13] a method for evaluating energy saving results has been proposed. The acceptable harmonic distortion levels in the vessel's network are discussed in [14] and solutions against line current harmonics produced by VFDs are presented in [15].
In this paper, a proposed Ship Energy Efficiency System (SEES) is developed and evaluated in order to improve E/R power balance and reduce DIG fuel consumption. The optimization of vessel's air ventilation and cooling water system involves variable frequency control of EIR Fans and Cooling Sea Water Pumps (C.s. W PIPs) according to temperature and pressure feedback signals from the process. The proposed solution builds on air balance and heat dissipation calculations for derated engine condition according to the vessel's operating profile. The parameter adjustment for the control algorithm depends on both performance and safety requirements. The improvement in energy efficiency is verified by case studies, simulations and experimental results following the application of the system onboard a 105,000 DWT tanker vessel. Onboard measurements show considerable energy and fuel savings, reduction of emISSIOn factors (CO/SO/NOx), while power balance improvement allows use of only one DIG in sea-going period, resulting in additional performance boost. Power quality measurements verity system compliance with marine harmonic standards. The system is designed and implemented as a retrofit solution for existing vessels according to classification society requirements.
In particular, Section II describes the limits established by PMS and compares vessel load balance during shop tests with current ship operating profile. In section III, the air ventilation and cooling water system energy saving margins for derated engine condition are calculated for a typical tanker vessel. The proposed process control scheme for the improvement of vessel's power balance is presented in Section IV. In section V, simulation and onboard experimental results are provided to verify the effectiveness of the control strategy and optimized DIG operating profile.
SLOW STEAMING OPERATION FOR A TYPICAL MARINE VESSEL
The PMS onboard monitors total power demand, compares it with the available supply and automatically starts and stops DIGs to coincide with load changes in accordance with pre-calculated startlstop tables. In case of sudden DIG failure, the load will be transferred to remaining DIGs online, which must take the rest of system load. The transient frequency deviation is limited to ±10% according to classification society rules. Failure to meet this limit leads to circuit breaker trip and consequently to blackout. Therefore, in a system where Nfunits rated at Pr,gi are tripped with k units online, the DIG load, P gi, is limited
by maximum continuous safe 10adPc:::'l/k,Nr), as in (1):
(I)
pr""X is determined by difference between transient load, cont,gi
P1rol1>g;{k,N;J, and maximum transient load step, as in (2):
where amax,g is maximum permitted engine load constant, which is usually between l.1 and l.IS. Pgi is also bounded by the recommended minimum engine load, as in (3):
0.15:::; amin,g :::;0.30 (3)
In order to prevent a blackout, the system must always have a sufficient power reserve or available power to its full online capacity. The minimum available power reserve, based on DIG failure cases is determined by (4):
The PMS also performs equal load sharing between DIGs according to classification society regulations, based on the load sharing constant Sdk). Equation (S) holds in order to avoid load limiting or reverse power condition of any DIG:
k k k P� = IS�/k)P� = IPg/k), where IS�/k) = 1 (S) i=1 i=1
Based on the above, the usual maximum permitted DIG load for a marine tanker vessel is 80% of P r,gi , while minimum load is IS% of P r,gi' DIG SFOC is the specific fuel oil consumption rate (g/kWh) calculated by (6):
SFOC = Mo(L) x C(0.91) x En (9750kcall kg)
(6) T(sec)x �,g/kW)x E\(10200kcall kg)
where Mo is the DIG fuel oil measurement (L), C is the marine oil specific gravity, En is the Heavy Fuel Oil (HFO) calorific value, T is the fuel oil gauge measurement time period (sec) and Es is the Fuel Oil Standard (MDO) calorific value. Equation (6) calculates SFOC for MDO, so calorific value correction for HFO has to be made if required. SFOC is a convex curve with minimum values at about 80% of Pr,g;' This means that, given an upper 80%
DIG load limit, SFOC is much higher when the same load is serviced by two parallel DIGs operating at low load than by only one DIG. Based on (6), the DIG Fuel Oil Consumption (FOC) in grams can be calculated as in (7):
FOC = p',gi_inCkW) x RT(hours) x SFOC (7)
where P r,gi in is the DIG mechanical input power and RT is the running time. The DIG emission factors depend on FOC and can be calculated, as shown in (8):
CO2(tn) = FOC(tn) x CFO = FOC(tn) x 3.17 SOo(tn) = FOC(tn) x S%x 0.02 NOx(tn) = FOC(tn) x 0.079
(8)
where CFO is the conversion coefficient of fuel oil and carbon emission for marine heavy duty diesel engines [16].
Nowadays, the use of slow steaming has changed the load balance onboard vessels in comparison with shop tests, especially during sea going period, having a negative impact on the DIG operating profile in terms of fuel consumption. Fig. I compares onboard load balance between sea trials and current vessel operation for a typical tanker at sea going, port inlout and cargo handling periods.
It is shown that at shop testlsea trials and according to original load balance study, power consumption is SI4kW or 6S.9% of the DIG capacity (Pr,gi=780kW) during sea going period. However, due to actual M/E operation at lower rpm (% of MCR) during sea going period, the intermittent load includes operation of both M/E Auxiliary Blowers (extra 1l0kW) to compensate for the required M/E scavenge air flow, which can not be provided by M/E turbochargers at this stage. This results in 624kW load or 80% of DIG capacity, which is not permitted by PMS due to (1). Consequently, PMS performs load sharing and now 2 DIGs are used in parallel, operating at 40 % of their nominal capacity, resulting in much higher SFOC.
In addition, since the starting current of most motors onboard can reach 8 or 3 times their nominal current value for D.O.L or Y//1 motor starting mode respectively, this has a considerable impact on voltage waveform at the Point of Common Coupling (PCC), as measurements show in Fig. 2.
Typical Aframax �/G Tanker Electrical 1 Distribution
DIG,
� D/GY
� O/G'=t1 3 Diesel Generators
3 (780kW 450V 60H ) ACB T ACB T ACB T z
l Point of Common Coupling (PCG) l Continuous Loads Intermittent Loads OJ OJ OJ OJ
5�W �W
e t_ e t_ '0 o e '0 o :J o e o :J 0)= 0= %' 0.0 �'" %' o.,Q �'"
- c '" e - e '" e Ujg « .- U '" Q) « .- Q) .<:: (f) (f) 444 681 876 70 63.5 NO.1 MIE NO.2 MIE 60 Aux. Blower Aux. Blower kW kW kW kW kW kW Additional loads used
(0.4 Group Div. Factor) continuously at sea going As per load balance As per load balance and at port in/out due to study at shop tests study at shop tests M/E slow steaming
Operational Profile Sea going (4500h) At port in/out (1800h) Cargo handling (2460h)
Power Consumption at 514kW, 65.9% of 744.5kW, 95.4% of 936kW, 120% of
Shop Test/Sea Trials 1st DIG capacity 151 DIG capacity 1st DIG capacity Need for 2nd DIG use? NO YES, at 47.7% YES, at60%
Power Consumption at 624kW, 80% of 854.SkW, 109.5% 936kW, 120% of current ship Oper. Profile 15t DIG capacity of 15t DIG capacity 151 DIG capacity
Need for 2m DIG use? YES, at 40% YES, at 54.7% YES, at 60%
Fig. 1. Onboard load balance comparison for a typical tanker
vessel between sea trials and current operation profile
L. M,""",ax L 12 RMS Vwgt (Cyclt by Cycle� ELSPEC12@ELSPEC-NTUA 24/05 25/05
Fig.2. RMS voltage measurements performed at 440V feeder
panel in Main Switchboard (MSB) for a 7-day period
As shown in Fig.2, load switching that often takes place during the operating period does not only lead to voltage sags but also requires use of 2nd DIG by PMS in order to satisfy the power reserve condition and prevent blackout.
III. E/R VENTILATION AND COOLING WATER SYSTEM
OPTIMIZATION POTENTIAL FOR DERATED ENGINE CONDITION
To improve ship's power balance and DIG operation especially during sea going mode, the energy saving margin through the optimization of EIR ventilation and cooling water system will be examined. Reduced flow capacities for derated engine condition can be satisfied if variable frequency control is applied to E/R ventilation fan and C.S.W PIP motors depending on process feedback.
For E/R ventilation fans, the power consumption (Pow) for a specific pitch angle, /3, air density, p, and impeller diameter, d, is determined by (9):
Pow = Pfi X Qoc px n3 x d5, where Qcxn (9)
The total fan impeller pressure (Pji) is the differential pressure developed between the inlet and outlet of the ventilation duct in a forced ventilation system as in a typical marine vessel, In EIR, Equation (10) must always hold for overpressure condition:
pft = Pin - Pout> 0 (10)
The fan performance curves variation through variable frequency control is compared with simple fan pitch angle adjustment and is shown in Fig. 3.
So far onboard, a total of 4 fans operate at full speed to supply adequate air flow for combustion needs through
Fan total Pressure (mmAq)
Pft1---+
Impeller pressure
change
P+ft2-
�f��� r r' r
�'>'� ,,-"'-
�"T r l-f-.
o
, 's,
iY"
;v >,
ved bv
, c c--
A c , "
,;>- [\ /"i ;$-
"'- ."
:i<"\L. /.
, i /1; OY ,
�, tF1J'J
� /'1 0
I l':s;; �
�; '\ / i i\
t::; K�� � G.> '''�T -/ �+ , ..... i""
, ; -. ; .-, -
"." ,.-. , .-'''-
"" '''-"" .
..... , -��
'''-i"-
..... , -""" '''-'i2 Flow q, Air Flovv (cmm)
- -I I change ,
Fig.3 : E/R Fan energy saving potential through frequency control
separate air ducts to M/E, DIG and Aux. Boiler, which are the major air consumers in E/R. The E/R Fans are designed at shop tests to operate at 100% of M/E MCR, so they are oversized with respect to current vessel operating profile. The flow requirements (m%) for the above consumers depend on operating % of MCR according to respective power outputs and air requirements as shown in (I I):
qM It = %PM It x 0.002?(kgl k�)/1.13(kgl m3) q lJiG = %PlJ!(, x O.OO2(kg 1 k�) x O.S 81 1.13(kg 1 m3)
m, (kg 1 s) x 0.0 77(kg 1 kg) x IS.7(kg 1 kg) x 0.911 (II)
qb = 1.13(kglm3)x3,600s1 h where PM/H is M/E output power, Pm; is DIG output power and m, is boiler steam capacity. The E/R Fans are oversized to 150% of required air flow at 100% of MCR:
qtot = 1.5 X �qM I r: _10Cl'1o + qTJ/G _1 0Cl'1o + qh _1 OCl'/J (12)
Required air flow for derated engine condition is calculated for a typical Aframax size tanker vessel and is presented in Table I, while the proposed E/R Fan variable frequency topology in order to match this demand is shown in Fig. 4.
TABLE 1: AIR BALANCE CALCULATION FOR AN AFRAMAX TANKER
Air balance calculation for derated engine required air flow
Operating condition Property
EJR Fan Power (kW)
EJRFan Flow (m'/h)
Technical Data DIG 100% MCR(kW)
M/EIOO% MCR
Boiler lOO%MCR
M/E operating %MCR
Value
II
50,400
730
11,488
25,000
At Sea-going and MIE required air flow, qM/E (m3/s)
54.9
12.84
40
0.6
at port in/out
condition D/Goperating %MCR
At loading and
unloading condition
Required air flow for
derated engine
condition
DIG required air flow, qDlG (m' Is)
DIG operating %MeR
DIG required air flow (m'/s)
Boiler working capacity (kg/h)
Boiler Required airtlow,qb (m]/s)
Max airtlow required, qtol (m3/s)
Number of fans required
60
0.6
22,000
5.96
20.16 10.28 at sea at port
1.44 0.73 at sea at port
------------ ,
E/R Vent. • Fan NO.1 - 4
Consumers that require air supply for combustion
Main Engine (M/E)
Boiler DIG o i �irfl�w �
: dlrectlon ______ � qb qDIG
Neutral room or outside area
, I:,
Ventilation duct area
I = Engine : qoul =1.5 x (qMfE +qDfG +qb) I ,Room area I i (E/R) � ___________ .J
Fig.4: Proposed variable frequency control for E/R Vent. Fans
Regarding the centrifugal cooling sea water pumps used onboard, power consumption (PBHF) is proportional to flow rate, q (gpm), total dynamic head, H (ft), and is related to specific gravity, Sg, and efficiency 1'/, as shown in (13):
HxqxS PRHP = 3,960 X 1]g
(13)
As shown in Fig.S, so far onboard throttle valves were used to reduce flow by increasing the slope of the system curve. This resulted in about 10% less energy consumption, but also in increased impeller pressure, vibration and noise. Through variable frequency control, however, the whole pump curve shifts downwards (point C), achieving the same reduction in flow q, but at significantly less pressure. The great energy saving potential through frequency control relies on affinity laws which express the relationship between motor speed (n) , water flow and consumed power:
:: = �: = � Z: = ti (14)
A simplified block diagram of a typical cooling water system for marine vessels is presented in Fig.6. In most cases, central cooling water system is used, where the dimensioning of heat exchanger (central cooler) was made for vessel's continuous operation near the nominal M/E rpm and MCR point (L/). Based on heat balance study, the heat dissipation of MIE scavenge air cooler, lubricating oil cooler and jacket water cooler determines the size of main
Total Dynamic Head (m)
Impeller
pressure
change
Flow q1 -,
change
WaterFlow (m'/h)
Fig. 5 : C .SW PIP energy saving potential via frequency control
High sea chest sea water inlet
,- ' Aux.
C.S.WPIP
Variable Wref Frequency 1------' Cooler
cooling sea water pumps. Due to slow steaming operation and current ship operating profile, the derated cooler capacities regarding heat dissipation, Q, at random point M (% of M/E MCR), where PM<PU, can be calculated by the following expressions, as shown in (15):
Qair,M = Qair,Ll X (Qair% /100) Qjw,M = Qjw,Ll X (QjW% /l00) (15) Qlub,M = Qlub,L1 X (Qlub% /100)
Central cooler heat dissipation at point M is given by (16):
Qcent"M = Qair,Ll + Qjw,M + Qlub,M (16)
Heat dissipation reduction factors can be found according to respective M/E model project guide.
The seawater flow capacity q (m%) for the central C.S.W PIPs can be reduced in proportion to the reduction of the total cooler heat dissipation as shown in (17):
q sW,cent,M = q sW,cent,Ll X Qcent,M / Qcent,Ll (17)
As jacket water cooler is connected in series with lub. oil cooler, the fresh water flow capacity for the latter is used also for the jacket water cooler. On sea water side of M/E central cooling system, water is supplied from two sea chests through the main C.S.W PIPs, which are currently running at full speed. An auxiliary sea water pump is also used during anchorage or in port. During sea going mode, S. W is transferred overboard through the atmospheric condenser, which can be by-passed. During unloading operation, vacuum condenser is in service to run the steam cargo pumps, using steam generated from Aux. Boiler.
Required sea water flow for derated engine condition is calculated for a typical Aframax size tanker vessel and is presented in Table II. Heat dissipation reduction factors refer to S60MC-C M/E model operating at 54.9% of MCR and can be calculated according to M/E project guide [17]. Variable frequency control of Main C.S.W PIP according to process feedback is proposed to adapt to derated flow requirement in order to benefit from energy saving potential.
Sea water system Fresh water system Jacket water system
control Qcent =
Qjw+Q/Ub+Qair
Fig.6 . Proposed variable frequency control for NO.1 C .S . W PIP, involving adaptation to reduced flow capacities for derated engine
TABLE IT: HEAT BALANCE CALCULATION FOR AN AFRAMAX TANKER
Heat balance calcnlation for derated engine reqnired water flow
Operating condition Property Value
M/Emodel S60MC-C
Technical Data Speed at 100'10 ofMCR (rpm) 97
Load at 100'/0 ofMCR(kW) 11,488
M/Eair cooler heat dissipation, QOi"Ll (kW) 5,800
lob, oil cooler heat dissipation,Qlob,Ll(kW) 1,130 Nominal capacities for
Jacket water cooler heat dissipation, QwLl(kW) 2,090 100'10 ofMIEMCR Central water cooler heat dissipation, Q"ot,Ll(kW) 9,020
Sea water flow capacity for central cooler qceot,LI (m'/h) 445
Speed at operating % of MCR (rpm) 79.5
Actual M/Eoperating Load at operating % ofMCR(kW) 6,318
profile (54,9% ofM/E Air cooler derating factor, Qoi�%) 43 MCR)
lob, oil cooler derating factor, Qh,b(%) 55
Jacket water cooler derating factor, QJw(%) 64
M/Eair cooler heat dissipation, Q"i"N(kW) 2,505
Lub, oil cooler heat dissipation,Qh,b,M(kW) 626 Derated capacities for
54,9% ofM/EMCR Jacket water cooler heat dissipation, QiwN(kW) 1,338
Central water cooler heat dissipation, Q"",,M (kW) 4,469
Sea water flow capacity for central cooler, q"otN (m'/h) 220.5
IV, DESIGN OF THE PROPOSED ENERGY EFFICIENCY SYSTEM
FOR A TYPICAL MARINE VESSEL
To match the derated air and water flow capacities as calculated above for the ship's operating profile, frequency control is used to regulate motor speed according to feedback from the process, The variable frequency drive topology as well as the design of the proposed process control algorithms will be presented in this section,
A, Proposed Variable Frequency Drive (VFD) topology
The proposed variable frequency topology, which is applied to motors in all processes, is presented in Fig,7, It comprises a 3-phase diode rectifier and a two-level 3-phase Voltage Source Inverter (VSJ) equipped with IGBTs, A 3% dc-choke is used in the dc-link in order to reduce harmonic distortion, An additional ac-choke is used in the c'S,W PIP case due to higher motor capacity in order to further reduce harmonic distortion, Sensorless vector control is used for switching process of inverter semiconductor switches,
The speed reference (OJref) to the VFD is provided by the output of the proposed Ship Energy Efficiency System (SEES) and the respective process control algorithms,
VF 0 Diode rectifier Ldc � ,__________ Three Phase VSI 3 phase motor
r-;-;--' ___ ._----.." Vdc c tt�H+:;::::::::'�
�is: Lac if! S, S, SS
� ' ----: I � �--������
Optional ae Choke
Fig.7: Proposed Variable Frequency Drive (VFD) topology
B. Proposed Process Control of E/R Air Ventilation System
The proposed topology for the control of air ventilation process is presented in Fig.8, while the respective control algorithm is presented in Fig.9. This configuration is valid for all 4 E/R fan motors onboard a typical tanker vesseL
E/R fan motor speed reference is provided in order to always maintain a slight overpressure in E/R as shown in (10), regardless of M/E, DIG and Aux. Boiler combustion air requirements at any given moment. EIR Fan speed is also regulated in order to ensure that E/R temperature does not exceed a certain set-point.
To achieve this, PID control provides a speed reference signal, OJref,ditfpres." in order to maintain 0.5mbar positive differential pressure between EIR and outside. Another PID controller provides a second speed reference signal, OJrej,TR.felllp' in order to maintain E/R temperature at 36°C. The fmal speed reference is the maximum value of the two signals as given by (18):
In order to avoid motor overheating due to very low rpm, the reference speed provided by the VFD is bounded between minimum and maximum limits, as shown in (\9).
O.55xwnom <Wref,Fan <O.95xwnom (19)
The lower limit can be set down to 55% of OJnom, because the E/R ventilation fans are self-cooled through the individual ventilation ducts.
Neutral room or outside area
SEES process
control
E/R temperature D·ff. pressure
<!:--i----__\. DP Pn= PE/R -Pout
Air flow -----. direction ===!
E/R Fan NO_1 - 4 Ventilation duct area
qout
Engine Room area
(E/R)
Fig. S: Proposed topology for air ventilation process control
Start Choose Local VFD or Remote VFD control (HMI)
Fig. 9: Proposed algorithm for control of air ventilation process
Several additional safety functions are implemented in the control algorithm to ensure that system complies with unmanned ship mode (UMS) as required by classification societies. In case of controller and sensor failure or communication error condition, the motor reference speed is set to 95% of Wnam, which is the adjustment of the local potentiometer, providing a second manual speed reference. The system also deactivates faulty VFDs and regulates the speed reference of the rest in order to keep a minimum of two VFDs working. In case of blackout recovery, the system starts automatically in its previous condition.
C Proposed Process Control o/Cooling Water System
The proposed topology for the optimization of vessel's central cooling system is presented in Fig.IO, while the respective control algorithm is presented in Fig.ll. The configuration is valid for one out of two C.S.W PIPs if each pump satisfies 100% of required water flow at MCR or for two out of three C.S. W PIPs if each pump satisfies 50% of required water flow at MCR.
In the old configuration, a thermostatic controlled 3-way v/v located at the central cooler F.W. outlet maintained 32°C for cooling water, by circulating Fresh Water (F. w.) either through the central cooler or through a by-pass route. Using this configuration, the main C.S.W PIPs operated at full speed along the vessel's operating cycle.
With the proposed topology, C.S.W PIP speed reference is provided to maintain constant F.W. temperature (Trw. ) in sea going and port in/out conditions for M/E, DIG and aux. machinery cooling and at the same time maintain constant discharge pressure on the common sea water line. For unloading operation a constant temperature difference is maintained between sea water outlet and inlet
to compensate for the use of vacuum (Co.P. T) condenser. In order to match the reduced water flow requirements
for derated engine operation while sea going and port in/out conditions, PID control is used to provide a speed reference signal, Wre!.lempFW, in order to maintain F.W temperature 2°C more (34()C) than the setting of the 3-way v/v (in this case 32°C). The lower setting of 3way-v/v old controller will allow earlier opening of the v/v and more water flow passing from central cooling, boosting the performance of the new temp. controller during transients. A C.S.W PIP discharge pressure PID controller provides another speed reference signal (wrej,s.w.press) in order to maintain S.W. pressure (Puvaul) at 1.2bar, allowing I m/s minimum water flow velocity through the piping system.
Start Choose Local VFD or Remote VFD control (HMI) or system by-pass (Y/!':,,)
Speed ref Wref 1 defined by -
temperature PI control, target
TFW= 34°C
Set Speed Reference Wref = max {wreU ,WreC2}
2
70% of �"'IL ::; Wref ::;95% of �"'IL
F ig.ll: Proposed algorithm for control of cooling water system
Condenser By-pass v/v
r - - --t:*::i-- - - - --, 3-way
thermostatic I tj Expansion tank
central cooling water
Sea water system
Fresh water system Jacket water system Temperature meas.
Sea Going Control t ,-----------c- + valve I :Atmosphenc ... I --: Condenser �---- - 1
� I S.Wout
1 $
F.Wout Central Cooler
S.Win F.Win q;n I Pump t Cooling I - - 1 � Outlet Fresh Water
_ Vacuum T pressure - -J>I<J-1 condenser 1- - I P,!mps
se:���er
q;-�::�r�
l
p �.�.� Cooling Cooling
Sea Wate Sea Water Pump
I Pumps . - - - - - 1-- .... ---4 _ � T}_ T
Inlet High sea chest sea water inlet Temp.� __ L-L-____ -.
3-way v/v
osition
------ Cargo Handling Control I-r--'-�.-------I>'<l---
$ Main Engine
I Lub. Oil cooler
to jacket cooling N' water system - �
$ Jacket Water
cooler
.... from F.W.
p
Scavenge air cooler
Air coo/er enerator /
, - - - [c' '-- , �---[7/: +---[1'-- :
SEES Process control
F.W. temp . 1'----. . 0( . I i----±=J1 I
t t t.. t t.. t � D/Gi � D/G2. I � D/G3
Low sea chest sea water inlet
to jacket t �_�J r � � _ _ J r � ��\ J
S.W. Outlet Temperature cooling water ... __ . ...1_._._._._._._._._._ ....... _._._._ _ ____ ._ . ...1 Engine system Lub. Oil cooler driven water
pump
Fig.lO. Proposed SEES topology for process control optimization of a typical marine vessel's central cooling water system
In order to avoid motor overheating due to very low rpm, the reference speed provided by the VFD is bounded between minimum and maximum limits, as shown in (21).
0.70 X OJnom ::;; OJrej,PP _seagoing ::;; 0.95 X OJnom (21)
Since the pump motor has no self-cooling mechanism, the lower limit value is set higher than in the E/R fan case.
While unloading, the control algorithm by-passes the F.W. temperature controller and differential temperature PID control is used to provide a reference signal, Wrefdifftemp in order to maintain a temperature difference between vacuum condenser sea water outlet (Tuvout) and sea water inlet (Tsw in) at 3Aoe. The temperature difference is created due to steam cargo pumps which are in service when unloading. To maintain a sea water pressure higher than the existing pressure switch limit (0. 7bar), the reference speed is bounded by values as shown in (22):
0.80 X OJnom ::;; OJrej,PP _unloading::;; 0.95 X OJnom (22)
The control system complies with the safety functions described before to ensure that classification society requirements regarding unmanned ship mode (UMS) are met. However, since this is a critical application, additional safety functions are added regarding fail-back capabilities. A by-pass mechanism is added in order to be able to revert to the old starter at any given moment, while the standby pump starts in case S.W. pressure drops below 0.7bar.
V. EVALUATION OF RESULTS FROM THE PROPOSED CONTROL
SCHEME IMPLEMENTATION ON BOARD A TANKER VESSEL
Following the results of case studies and the principles involved, the proposed process control schemes are integrated into a Ship Energy Efficiency System (SEES) and are applied onboard a typical I05,000DWT Aframax size tanker vessel, which has a M/E operating profile similar to the one described above. The proposed frequency control topology is applied to four IlkW, 440V, 22.3A, 60Hz E/R ventilation fans and to one of the two 75kW, 440V, 129.IA, 60Hz main e.s.w PIPs. In this section, the experimental results regarding the efficiency of the proposed control scheme are evaluated, a power quality analysis of the proposed solution is provided and DIG optimized operation is verified after the system application.
A. Experimental results {rom the optimization o{vessel's air ventilation system
The respective outputs of the E/R ventilation fan controllers are shown in Fig.12. A comparison of the starting current waveforms between the old D.O.L starting method and the proposed frequency control is presented in Fig.I3. The voltage and current control waveforms applied to the motor from the inverter as well as the equivalent grid side voltage and current supply waveforms are shown m
Fig.I4. NO.1 E/R Fan active power consumption IS
presented in Fig. 15 after a IO-day logging period.
(a) (b)
Fig.12: E/R ventilation fans control scheme
(M/E operating at 54 .9% of MeR - 6,318kW, 4 Fans working)
(a) Diff. Press. PID control (P:2. 1:20. D:I. Min:55%. Max:95%)
(b) EIR Temp. PID control (P:I. 1:2. D:2. Min:55%. Max:95%)
Inrush Current O!ill 2013-04-18.20:46 Trace _ . Apk +171.1,-172.7
.. A ....... ................................................... +300 pk····· h
E ....... ................................................... -300 T - 1.25 10 sec
(a)
Inrush Current ill} 2013-05-11. 14:30 Trace _ . A pk +31.31,-31.24
As shown in Fig. 12, the PID differential pressure controller speed reference is 66.S% in order to reach the set-point of 0.45mbar, while the respective E/R temperature PID controller speed reference is the lower limit of 55% as in (19) , since actual E/R temperature (30.9()C) is much lower than the set-point value of 36°C. Therefore, for this case the differential pressure controller output determines the common speed reference for the E/R fans according to (IS). This results in around 3.0 kW power consumption for each fan, which is much lower than the original llkW. As shown in Fig. 13, there is also considerable improvement regarding the reduction of peak starting current values, since the original D.O.L starter caused a current peak of 171 A in comparison to the gradual current raise to 22.3A full load current using the proposed VFD starter. From Fig. 14(a), it can be seen that the motor impedance acts as a filter for the applied current waveforms. The use of dcchoke partially compensates for the distortion to the supply current waveform, which is caused by the use of the 3-phase diode rectifier, as shown in Fig. 14(b). The power quality analysis presented later in this section, will verify that voltage harmonic distortion is in line with marine harmonic standards [14].
The 10-day period active power logging for NO.1 E/R fan which is presented hereby, confirms the effectiveness of the proposed control scheme. For the first five days, 4 E/R fans were in service and it is shown that individual E/R fan power consumption varied from 2,5kW to 5,5kW according to MIE, D/G and aux. machinery combustion air requirements. For the next five days, 3 E/R fans were in service due to NO.2 E/R fan maintenance, so the controller raised the speed of remaining fans to maintain set-points.
Based on onboard experimental results, the energy saving study is performed and presented in Table III, while the results are presented in Fig. 16.
TABLE Ill: SEES ENERGY SAVING ANALYSIS FOR E/R FANS
Sea going + unloading + loading + port in/out anchorage
51.90% 25.95% 22.15%
4 3 2
II II II
65% 70% 75%
3.02 3.77 4.64
12.08 11.32 9.28
44.00 33.00 22.00
31.92 21.68 12.72
4,546 2,273 1,940
145,100 49,286 24,700
22,400 7,600 4,300
2,000
36,300
350 300
� O!: 250 .� 200 [0150 � 100
50
Year profit marlin in MWh from VFO application to E/R Fans (Aframax) tanker
219 275 1-------------1 318
18
Sea going-unloading-port
In/out
loading/anchorage Total Energy consumption
E/R VlI!!ntilation Fans Enll!!rgy consumption in lI!!ach oplI!!l'ating mod II!!
• Energy Consumption before SEES application • Energy consumption after SEES application
Fig.16: Energy saving margin for E/R fans from SEES application
Significant fuel and energy savings are calculated from the proposed system application. The study is performed based on 630$/tn HFO price, resulting in 0.152$/kWh cost and on 900$/tn MOO price, resulting in 0.20$/kWh Cost. MOO is mostly used while unloading or in port mode owing to environmental regulations.
B. Experimental resultsfrom the optimization of vessel's central cooling water system
The outputs of the NO.1 C.S.W PIP PID controllers in seagoing and port in/out modes are shown in Fig.17. Fig. IS shows the outputs of the PID controllers used in unloading mode. A comparison of the starting current waveforms between the old Star/Delta starting method and the proposed frequency control is presented in Fig.19. The voltage and current control waveforms applied to the motor from the inverter as well as the equivalent grid side voltage and current supply waveforms are shown in Fig. 20. NO.1 C.S.W PIP active power consumption is presented in Fig. 21 after a 10-day logging period, showing the difference in consumption before and after the proposed frequency control for energy saving optimization.
(a) (b)
Fig.17: NO.1 C .S .W PIP sea going-port inlout control scheme
(MIE operating at 54 .9% of MCR - 6,31SkW, NO.1 PIP working)
(a) F.W. Temp. PID control (P:1O,1:100,D:l,Min:70%,Max:95%)
Fig.21: NO.1 C .S .W PIP active power (P) logging in kW for 10
days period after the proposed control scheme implementation
As shown in Fig. 17 for seagoing, port inlout and anchorage condition, the F.W. Temperature PID controller speed reference is 70% in order to reach the set-point of 34°C, while the respective S.W. Discharge Pressure PID controller output is 70.2% in order to maintain the set-point of 1.2bar discharger pressure on the common sea water line. Therefore 70.2% speed reference is applied to the motor, since this value satisfies both (20) and (21). The setpoints are chosen in such way so that balance is kept between energy savings and ship reliable operation. In Fig. IS, it is shown that the above controllers are by-passed in unloading mode, since M/E is not working and the differential S.W. temperature controller is used instead, providing SO. 1 % speed reference in order to keep the temperature difference between vacuum condenser S.W outlet and common S.W. inlet at 3AoC. This speed reference is applied to the S.W PIP as it complies with (22). Therefore, it is shown that in sea-going mode, the pmnp is running at 70.2% speed, resulting in 25.SkW power consumption which is much lower in comparison to the
nominal 75kW. In unloading mode, the pump is running at SO.3% speed, resulting in 39kW power consumption, creating a profit margin of 66kW. This is because both the NO.1 C.S. W PIP and the Auxiliary C.S. W PIP were used at unloading before, the former for the vacuum condenser and the latter for vessel 's cooling through vlv isolation. With the new controller, NO.1 C.S.W PIP alone can service both systems while running at reduced speed.
As shown in Fig.19, there is also reduction of peak current values during starting operation, since the original StarlDelta starter caused much greater current peaks in comparison to the gradual current raise to 129.1 A full load current using the proposed VFD starter. From Fig. 20(a), it can be seen that the motor impedance acts as a filter for the applied current waveforms. The combined use of dc-choke and ac-choke further compensates for the distortion to the supply current waveform which is caused by the use of the 3-phase diode rectifier, as shown in Fig. 20(b). The power quality analysis will verify that voltage harmonic distortion complies with harmonic standards [14].
The 10-day period active power logging for NO.1 C.S.W PIP, which is presented hereby in Fig.21, verifies the effectiveness of the proposed control scheme. The proposed frequency control is applied first time between 13/05 and 14/05. It can be easily seen that 350kW are required from DIG during starting in original StarlDelta mode, causing potential need for running an additional DIG in parallel. This demand is normalized down to the nominal power with the proposed SEES application. It is shown that NO.1 C.S.W PIP frequency control results in varying power consumption from 24kW up to 60kW according to MIE and DIG operating profile. The power factor (coscp) is also improved from O.S to 0.95 at all times.
Based on onboard experimental results, the energy saving study is performed and presented in Table IV, while the results are presented in Fig. 22.
TABLE TV: SEES ENERGY SAVING ANALYSIS FOR NO.1 C .S .W PIP
Cooling Se a Wate l' Pumps e n e rgy saving analysis
Year Profit margin in Mwh from VFD application to c.s.w pIp (Aframax)-2xlOO% 700
� 600 ;: 500 � 400 i:i 300 � 200
w 100 0 -'---- -.. 54 20
Sea going - port in/out loading/anchorage U n l oading Total Energy co n s u m pt ion
(.s.w pIp Energy consumption in each operating mode • Energy consumption before SEES appl ication • Energy consumption after SEES application
Fig.22: Energy saving margin for S . W.P/P from SEES application
C. Power quality measurements after SEES application
In Figs. I 4(b) and 20(b), it was shown that since the variable frequency drives use a 3-phase full-wave diode rectifier, 5th ,ih and 11 th current harmonics are generated. These harmonics can cause considerable voltage distortion at the Point of Common Coupling (PCC) if left untreated, because the respective motors are constantly in operation. In this solution, a 3% dc-choke was used for the E/R VFDs, resulting in 43.3% current Total Harmonic Distortion (THDi%), as shown in Fig.23 (a). For NO.1 C.S.W PIP VFD, both 3% ac and 3% dc-chokes were used, to compensate for the higher level of load current. As a result, THDi was lowered to 22.9% as shown in Fig. 23(b). In both cases, the highest harmonic frequency was the 5th one with a value of 15% compared to fundamental frequency.
According to marine harmonics standards [14], the voltage Total Harmonic Distortion (THD,,%) should be less than 5% and no invidual voltage harmonic should exceed 3% compared to the fundamental frequency. In Fig. 24, it is shown that according to measurements performed when all VFD starters are in operation, the system is fully compliant with classification society requirements, as THDv% is 1.8%.
A A 30 % H 1 1 0
1 9 24 � 0 4
(a)
..
I I • 9 1 4
(b)
1 9 2 4
3 0 % H 1 1 0
Fig.23 : Power quality measurements of supply current waveform
(a) Harmonic distortion of NO. ! EIR fan VFD supply current
(b) Harmonic distortion of NO. ! C .S .W PIP VFD supply current
Fig.24: Power quality measurements of the voltage supply
waveform at the Point of Common Coupling (PCC) in MSB
D. Diesel Generator operating profile optimization results
The proposed process control schemes are integrated into a Ship Energy Efficiency System (SEES) which is designed to improve DIG operating profile. It was shown that the proposed system not only eliminates current spikes during motor starting but also creates a power saving margin of around 80kW during sea going period, which enables use of only one DIG, resulting in lower SFOC.
Following the analysis of Fig. 1, Fig.25 compares the variation of energy cost and DIG fuel consumption before and after the proposed system implementation for seagoing and unloading periods based on DIG SFOC curve. Before SEES, two DIGs operated in parallel at 40% of MCR during sea going, resulting in 210g/kWh SFOC, 0.20$/kWh energy cost and 3. 15tnlday HFO consumption. With SEES, the 80kW power saving margin allows use of only one DIG which operates around 70% of MCR. This results in 192.3g/kWh SFOC, 0.17$/kWh energy cost and 2.53tnlday HFO consumption. Considering that nearly 50% of the operating cycle involves sea going operation, this practice generates significant savings just from the improvement of DIG operating profile.
Fuel savings generated from the proposed system have a good environmental impact as well, since CO2, S02 and NOx emission factors are greatly reduced. It is shown that the proposed solution results in 750-800 tons HFO consumption during sea going days, well below the original nearly 1000tons consumption. Emission factors are calculated based on (8) for 2% sulfur content in HFO and are presented in Fig. 26.
Aframax Tanker Diesel Generator (DIG) operational profile optimization from SEES application
. : _$MWh at unloacing wout SEES I �$MWh at unloacing I'.ItIl SEES 0.20 ! _ DIG SFOC CU�.I€
I DIG upper load I mit as
I f3djusted from ower
I J. •• Management S stem
0 . 1 9
0 . 1 8
0 . 1 7
0 . 1 6 I I I I il l l l l i l l I I 1 8� 30 35 40 45 50 55 60 65 70 75 85 90 95 1 00°. 1 5 % of DIG load (bottom) and respective partial power output for balanced DIG in kW (top)
Fig.25: Power balance optimization resulting in improved Diesel
Generator operating profile and less fuel consumption
� Total Aframax Tanker DIG emission fa::tors variation at sea going days from SEES application � � 3500 : OAO � g : Emission Emission Fact g> "& : factors without SEE "& CIJ
: rth SEES (2 DIGs in use (IJ � I( DIG in use) � 1il 3000 0.35 : � c g � c 0 .� � E •
In this paper, an integrated Ship Energy Efficiency System (SEES) has been proposed and evaluated in order to improve E/R power balance and reduce DIG fuel consumption. Air and heat balance case studies were performed and showed that significant energy saving margin exists due to the vessel's current operating profile and slow steaming practice. To exploit the potential benefits, the proposed system involved the optimization of vessel's air ventilation and central cooling system through feedback from the process. The proposed control algorithm was implemented onboard a typical tanker vessel and managed to reduce total power consumption of EIR Fans and C.S.W PIPs through frequency control by 80kW during sea going period and by lOOk W during unloading period. Especially during sea going period, this power saving margin was verified to allow use of only one DIG, resulting in 0.62tons/day less HFO consumption. Maintenance costs were also reduced due to less mechanical and electrical stresses to DIGs and auxiliary machinery. The proposed system complies with classification society rules and regulations regarding operation in case of faults. Power quality analysis shows also full compliance with marine harmonic standards. The energy saving analysis which is performed based on onboard experimental results shows that the maritime industry can be greatly favored from this application which is proposed as a practical and effective retrofit for existing marine vessels.
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[ I ] P. Mitra and G.K. Venayagamoorthy: "Implementation of an Intelligent Reconfiguration Algorithm for an Electric Ship' s Power System", IEEE Trans. Ind. Applications, vol. 47, no.5, 201 1 , pp. 2292-2300
[2] Y. Durmusoglou, T. Satir, C. Deniz and A Kilic: "A Novel Energy Saving and Power Production System Performance Analysis in Marine Power Plant Using Waste Heat'", in Proc. 2009 IEEE Machine Le arning and Applications ICMLA '09 Con!, pp.7 5 1 -754
[3] C.L.Su, M.C.Lin and C.H.Liao: "A energy-savings evaluation method to justify automatic power factor compensators on marine vessels", in Proc. 2012 IEEE Industry Applications Society Annual Mee ting, pp. 1 - 1 0
[4] l Meyer, R. Stalbock, S.Voss: "Slow Steaming in Container Shipping'", in Proc. 2012 IEEE System Science HlCSS Con!, 2012, pp. 1 3 06- 1 3 1 4
[5] lA Rooks and A K Wallace: "Energy efficiency of VSDs'", IEEE Trans. Ind. Applications, vol. I 0, no.3 , 2004, pp.57-6 1
[6] E.P Wiechmann, P. Aqueveque , R. Burgos and J. Rodriguez: "On the efficiency of voltage source and current source inverters for high-power drives" , IEEE Trans. Ind. Electronics, vol. 55 , no. 4, 2008, pp. I 7 7 1 - 1 782
[7] l Arribas and C. Gonzalez "Optimal vector control of pumping and ventilation induction motor drives" , IEEE Trans. Ind. Electronics, vol. 49, no. 4, 2002, pp.889 -895
[8] Ade Almeida, F. Ferreira and D. Both: "Technical and economical considerations in the application of variable-speed drives with electric motor systems" , IEEE Trans. Ind. Applications, vol. 4 1 , no. 1 , 2005, pp. 1 8 8 - 1 99
[9] P. Kini , R. Bansal and R. Aithal: "Performance analysis of centrifugal pumps subjected to voltage variation and unbalance", IEEE Trans. Ind. Electronics, vol. 55 , no. 2, 2008,pp.562-569
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[ 1 1 ] F.lT.E Ferreira, lAC Fong and AT de Almeida: "Eco-analysis of Variable Speed Drives for Flow Regulation in Pumping Systems", IEEE Trans. Ind. Electronics, vol. 58, no. 6, 20 1 1 , pp. 2 1 1 7 -2 1 2 5
[ 1 2] Chu-Lien S u and Kuen-Tyng Yu, "Evaluation o f Diflerential Pressure Setpoint of Chilled Water Pumps in Clean Room HVAC Systems for Energy Savings in High-Tech Industries", IEEE Trans. Ind. Applications, vo1.49, no.3 , 20 1 3 , pp. 1 0 1 5 - 1 022.
[ 1 3 ] D.E Rice: "A suggested energy - savings evaluation method for AC adjustable-speed drive applications'", IEEE Trans. Ind.Applications, vol. 24, no.6, 1988 , pp. I I 07- 1 1 1 7
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[ 1 6] A Sarvi, C.J. Fogelholm, R. Zevenhoven: "Emissions from largescale medium-speed diesel engines: 2. Influence of fuel type and operating mode'", Journal of Fuel Processing Technology, Vol. 89, No. 5 , 2008, pp.520-527
[ 1 7] Man B&W Diesel and Turbo, "S60MC-C8 Project Guide Camshaft Controlled Two-Stroke Engines" , I sted., 2009, Man B&W, [Online ].Available:http://www.mandieselturbo.comldownload/proje ct_guides _tier I Iprinted/s60mcc8. pdf
A cascade control scheme for a grid connected Battery Energy Storage System (BESS)
Spyridon V. Giannoutsos, Student Member, IEEE and Stefanos N. Manias, Fellow, IEEE *
Abstract - The development of a Battery Energy Storage System (BESS) is considered to be very important for peak shaving and power demand normalization, especially in an autonomous power system. This paper presents a control scheme based on a cascade of an external controller that regulates the active and reactive power to generate reference currents and an internal current controller which produces gating signals for the inverter semiconductor switches. The proposed control algorithm takes the state of charge (SOC%) of the battery system into account and implements an active power command override to satisfy the requirements of the SOC% limits of the batteries. System simulations are performed during discharging and charging modes of operation as well as during the active power command override mode for several power demand step variations. The evaluation of the results shows good system performance and dynamic behavior. Index Terms - Battery Energy Storage System (BESS), P-Q control, State of charge (SOC), VSI grid interconnection
I. INTRODUCTION A promising method in order to regulate power
demand in a power system is to store energy when the demand is decreased and inject energy back to the grid at peak load. This can be realized through a battery energy storage system with a four-quadrant grid-synchronized inverter operation, which can store energy from the grid and inject energy back to the grid at a given active (P) and reactive (Q) power command. When a current control is used, the inverter output currents are measured and compared to reference signals in order to provide the voltage reference to an SPWM modulator which provides the gating signals for the inverter semiconductor switches.
In the past, several methods have been proposed for the control of such inverters [1]-[5]. The strategy presented in [1] focuses on islanding detection and fault limit, in [2] effective communication between DG Units through droop control and average power control is implemented, while in [3] unbalanced micro-grids are tested for inverter stability. A comparative study of VSI CSI as an interface for distributed generation (DG) applications has been performed in [4]-[5], while in [6] DG control schemes based on fuel cell, photovoltaic, and wind turbines are presented. Particularly for BESS control
*S.V. Giannoutsos and S.N. Manias are with the School of Electrical and Computer Engineering, National Technical University of Athens (NTUA), 9 Iroon Polytechniou st., 15780 Athens, Greece (e-mail: [email protected])
applications, [7]-[8] focus on the optimal operation of the system using linear and dynamic programming techniques, while control techniques for power conditioning and load leveling are presented in [9] .
The control scheme evaluated in this paper uses the state of charge (SOC %) of the batteries to satisfy the system limits, while regulating the active and reactive power through two control loops in cascade. The system uses a PLL to synchronize the inverter with the grid which makes it insensitive to disturbances, such as voltage sags. Simulations are carried out in MATLAB during the system’s battery discharging and charging modes for load step variations. An active power command override function is included, where the system does not follow the active power command in order to satisfy the SOC limits.
II. BESS TOPOLOGY AND MODES OF OPERATION To achieve peak shaving and load leveling, BESS
should charge the batteries with energy produced from base load power plants during low power demand and discharge at peak load, delivering power back to the grid to normalize power supplied by high cost power plants such as wind or photovoltaic power stations. During intermediate power demand, BESS is used to balance supply and demand in order to maintain constant frequency if it operates as part of an autonomous power system, as shown in Fig.1. During this operation, the system must satisfy the battery SOC limits.
Fig.1. Optimal charging and discharging of BESS batteries for peak shaving and load leveling, achieving max. economic benefits
The under examination BESS topology is presented
in Fig.2. It consists of a Li-ion battery pack, the four-quadrant three phase VSI, which is able to transfer power in both directions, an LCL output filter and the necessary voltage and current measuring equipment placed at the
2nd IEEE ENERGYCON Conference & Exhibition, 2012 / Future Energy Grids and Systems Symp
Point of Common Coupling (PCC), which provide inputs to the microcontroller-DSP that produces gating signals for the inverter semiconductor switches.
Fig.2. Block diagram of the under examination BESS topology
III. PROPOSED CONTROL SCHEME PRESENTATION Α. Current control using P-Q regulation
The block diagram of the proposed control scheme for BESS is presented in Fig.3. The P-Q control external feedback loop regulates the calculated active and reactive power using the following expressions: ccbbaa ivivivP ⋅+⋅+⋅= (1)
( )ccabbcaab ivivivQ ⋅+⋅+⋅−=3
1 (2)
The values obtained are compared to the reference P*, Q* commands to produce an error that is an input to PI controllers that produce the i*d_P and i*q reference currents respectively. A i*d_SOC reference current is also produced by another PI controller that compares the mean value of SOC levels at that instant with a suitable SOC limit. The resulting i*d depends on the power flow direction command at that time and can also be calculated from the P, Q setpoints defined by the following expression:
22
***
32
qd
qdd vv
vQvPi
++
= , 22
***
32
qd
qdq vv
vQvPi
+−
=
(3)
A positive P-command means that there is need for battery discharging, while a negative P-command means that the system absorbs power from the grid to charge its batteries. The grid current iabc is measured at the PCC and after it is transformed to a dq0 reference frame, rotating at the electrical frequency ωe of the grid voltage, it is inserted to a PI current controller that produces the reference voltages v*d and v*q ,which are transformed to abc frame using Park’s reverse transformation. The synchronization of the reference frame to the grid voltage is achieved through a suitable PLL method. Then, the gating signals for the inverter switches are produced using the SPWM technique.
B. Current control using battery SOC% levels The state of charge (SOC) for a battery is dependent
on the maximum battery capacity (Q) and can be described by the following expression:
⎟⎟⎠
⎞⎜⎜⎝
⎛−= ∫
t
dttiQ
SOC0
)(11100 (4)
BESS control must be able to keep the SOC levels in such values, depending on the predicted power demand, in order to provide power by discharging the batteries not on the specific instant the active power command is given, but during a load peak in the future. In order to achieve this, current control using battery SOC% levels must be carried out. The control algorithm, presented in Fig .4, accepts the SOC_upper limit and SOC_lower limit as input values as well as the battery SOC% levels at that instant. There is also estimation of the active power flow direction. If the SOC limits are not violated, then i*d = i*d_P. If the SOC limits are violated and the active power command has not changed its flow direction, then the active power command override control is activated and i*d = i*d_SOC.
( )
ccbbaa
ccabbcaab
ivivivP
ivivivQ
⋅+⋅+⋅=
⋅+⋅+⋅−=3
1
Nn1 SOC...SOC ++
Fig.3. Block diagram of the proposed BESS control scheme
470
Fig.4. Flow diagram of active power command override control using battery SOC% levels IV. SIMULATION AND PERFORMANCE EVALUATION OF THE CONTROL SCHEME
The BESS topology and the proposed control scheme are simulated in MATLAB for step load variations in battery discharging and charging modes. The battery active power command override control is simulated as well for two different cases.
BESS is able to handle nominal active power of Pnom = 1MW, while the maximum inverter dc voltage is Vdcmax = 1000V. The Idcmax = 2400A. BESS is connected to a 400V bus, 50Hz and the inverter switching frequency is 1050Hz, which is an odd number multiple of three. The 400V bus is connected through a transformer to a 20kV bus, where the grid short circuit power is o
k MVAS 8020~ ∠= . There is a P = 3MW load connected to the dc bus with cosφ = 0.85. A pack of two Li-ion batteries is considered with nominal dc voltage of Vdc = 650V.
Α. BESS discharging mode- delivering power to the grid
In this simulation scenario, the grid demands active power to satisfy load requirements. There is an active power step load change from 200kW to 900kW at 3sec and a reactive power step load change from 400kVAr to -400kVar at 7sec. The initial SOC levels are 85%. The upper and lower SOC limits are 90% and 30% respectively. Simulation results are presented in Fig.5.
(a) Active power command and response
(b) Reactive power command and response
(c) Inverter and grid voltage
(d) Grid current
(e) Reference voltage detail by the control algorithm
(f) Battery SOC% level
(g) Battery dc voltage
(h) Battery dc current
Fig.5. (a)-(h): BESS discharging mode waveforms
471
The battery discharge characteristics are presented in Fig.6. From the discharge curve and the nominal operating area it is observed that 2.3h are needed for the batteries to discharge completely under nominal discharge current.
Fig.6. Li-ion batteries discharge characteristics
B. BESS charging mode-absorbing power from the grid
In this simulation scenario, BESS stores active power due to low power demand or inexpensive energy production. There is a active power step load change from -300kW to -800kW at 3sec and a reactive power step load change from 100kVAr to 500kVar at 7sec. The initial SOC levels are 80%. The upper and lower SOC% limits are 90% and 30% respectively. Simulation results are presented in Fig. 7.
(a) Active power command and response
(b) Reactive power command and response
(c) Inverter and grid voltage
(d) Grid current
(e) Reference voltage detail by the control algorithm
(f) Battery dc voltage
(g) Battery dc current
(h) Battery SOC% Level
Fig.7. (a)-(h): BESS charging mode waveforms
472
According to simulation results, it is shown that active and reactive power responses need less than 1s until they reach steady state, while they display minimum overshoots. The switching frequency is fixed at carrier frequency because of SPWM technique. During the battery discharging mode (Fig.5), inverter phase voltage leads in relation to the grid voltage, which means that power transfers from the inverter to the grid. The battery SOC level is slowly decreasing due to discharge, while dc voltage is slowly decreasing as well. In this mode, the voltage source converter operates as an inverter, delivering power to the grid. The opposite occur in the battery charging mode (Fig.7), where the converter operates as a rectifier, charging the batteries. In Fig.8, the frequency spectrum of the grid current is presented and the THDi of 2.91% satisfies IEEE 519 and IEC 61000-3-6 limits [10].
Fig.8 Grid current waveform and its respective frequency
spectrum (THDi = 2.91%)
C. BESS simulation during active power command override Case 1: In this simulation scenario, BESS does not follow the active power command (P*) in order to satisfy the SOC limits.The system can return to the original state only if after a time period the SOC limits are satisfied and the active power flow command has changed direction. In this case, there is an active power command constant at 800kW and a reactive power step load change from 200kVA to -200kVA at 7sec. The upper and lower SOC limits are 90% and 89.998% in order not to discharge the battery now, but in the future. Simulation results are presented in Fig.9.
(a) Active power command and response
(b) Reference voltage detail by the control algorithm
(c) Grid current
(d) Battery dc voltage
(e) Battery dc current
(f) Battery SOC% level
Fig.9. (a)-(f): BESS active power command override control
(simulation of case 1) In Fig. 9 it is shown that even if the system initially follows the P-Q commands, when it reaches the 89,998 lower limit, the active power command override activates
473
and the desired target is now the mean value of the two SOC% levels. Thus, the battery begins to charge, which is observed by the negative dc current. The SOC% levels change periodically. Case 2: In this case, there is an active power step load change command from -800kW to 300kW at 11sec and a reactive power step load change from 100kVA to -300kVA at 7sec. The upper and lower SOC limits are 90.05% and 30% in order not to charge the battery further due to high energy cost .The initial SOC levels are 90%. Simulation results are presented in Fig.10.
(a) Active power command and response
(b) Battery dc current
(c) Battery SOC% level
Fig.10. (a)-(c): BESS active power command override control
(simulation of case 2) In Fig. 9 it is shown that the system initially follows
the P-Q commands until 9sec. At that moment it reaches the high limit of 90.05% the active power command activates. Thus, the battery begins to discharge delivering to the grid the maximum power that the inverter can handle. This control is valid until 11sec. At that moment, the active power command changes its flow direction and the system is within SOC limits. Therefore, from 11sec and after, the system follows the P-Q commands.
V. CONCLUSIONS
In this paper, a control scheme for a battery energy storage system has been presented and analyzed. The proposed control is realized by using two control loops in cascade. The external P-Q control loop produces the reference currents and the internal current loop regulates the gating signals for the inverter semiconductor switches. A Battery SOC% level control is integrated in the control scheme in order to achieve maximum efficiency in peak shaving and load leveling. The active power command is followed only if the battery SOC limits are satisfied. The simulations show that system responses are quick and display minimum overshoot. The power quality is within IEEE 519 and IEC 61000-3-6 limits as the THDi is well below 5%. Last, the switching frequency is fixed at carrier frequency, which is important for the power inverter semiconductor switches.
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