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A quantitative comparison of hybrid diesel-electric
and gas-turbine-electric propulsion for future frigates
Rinze Geertsma∗‡ , Jasper Vollbrandt† , Rudy Negenborn‡ , Klaas Visser‡ and Hans Hopman‡
∗ Faculty of Military Sciences, Netherlands Defence Academy, The Netherlands. Email: [email protected]
† Defence Materiel Organisation, Ministry of Defence, The Netherlands.
‡ Dept. of Maritime & Transport Technology, Delft University of Technology, The Netherlands.
Abstract—Future frigates need to reduce fuel consumption
and emissions, while improving effectiveness. This paper quantitatively compares top speed, fuel consumption, emissions, acceleration performance and engine loading of hybrid diesel-electric
and gas-turbine-electric propulsion, using validated models and
introducing sequential turbocharging and power take-off. This
simulation study demonstrates hybrid diesel-electric propulsion
can reduce fuel consumption and CO2 emissions by 10% to 25%
compared to gas-turbine-electric propulsion, while reducing top
speed by 3 kts. Moreover, hybrid diesel-electric propulsion is
found to provide good acceleration without thermally overloading
the engine when speed control is employed for the electric drive
in combination with torque control for the diesel engine.
Fig. 1: Artist impression of notional future frigate for the Royal
Netherlands Navy, case study in this paper.
I.
I NTRODUCTION
The Royal Netherlands Navy (RNLN) is under great political and societal pressure to reduce environmental impact
and logistical dependency on fuels. In the Operational Energy
Strategy, the Netherlands Ministry of Defence defines the
target to reduce its dependency on fossil fuels by 20% in 2030
[1], [2]. Therefore, future frigates need to reduce their fuel
consumption, while maintaining their effectiveness. Castles
and Bendre [3] have demonstrated the economic benefits
of hybrid propulsion for gas-turbine driven ships with shipservice turbine generators (SSTGs). Thus, many European
Navies use hybrid gas-turbine electric propulsion with diesel
generators, for example in the French-Italian FREMM frigates
[4]. These navies might further reduce their fuel consumption
with hybrid diesel-electric propulsion, because diesel engines
have a significantly lower fuel consumption than gas-turbines.
modelling strategy proposed in [9]. With this modelling strategy, we model a sequentially turbocharged diesel engine based
on project guide parameters and Factory Acceptance Test
measurements with the procedure proposed in [9]. Secondly, in
Section III we propose a parallel control strategy for propulsion
that uses the induction machine as a generator to provide power
to the auxiliary loads. This control strategy is used during
transit mode, when the ship is not required to accelerate. In
manoeuvre mode the control strategies proposed in [10] will
be used to investigate the manoeuvrability of the case study
frigate. After introducing these novelties, the proposed models
and control strategies are used to assess the performance of
a case study frigate of 5200 tons. We assess top speed, fuel
consumption, emissions, acceleration performance and engine
loading, with the performance criteria proposed in [9].
However, the limited power density of diesel engines will
either lead to an increase in ship weight or a decrease in
top speed. Moreover, the manoeuvrability of frigates with
diesel mechanical propulsion has been limited to prevent
thermal overloading [5], [6]. Using the electric drive and
controllable pitch propeller in a parallel hybrid propulsion
configuration could improve manoeuvrability of hybrid dieselelectric propulsion and increase top speed [7]. The subsequent
trade-off between acceleration, fuel consumption, emissions
and thermal loading can only be investigated adequately with
dynamic simulation models as proposed in [8]. Dynamic
simulation models have been developed and validated with
measurements on the RNLN Holland class patrol vessels [9].
II.
This paper compares hybrid gas-turbine-electric propulsion
with hybrid diesel-electric propulsion with advanced parallel
control strategies. The hybrid propulsion layout is shown in
Figure 2 for the diesel-electric variant. In the gas-turbine
electric variant, a gas-turbine replaces the diesel engine.
A. Hybrid diesel-electric propulsion
In order to investigate top speed, fuel consumption, emission, acceleration performance and engine thermal loading,
dynamic models are required [9]. We use a mean value first
principle approach to compare different propulsion architectures and control strategies, according to the modelling structure in Figure 3. The direction of the arrows shows causality
of coupled effort and flow variables as proposed in Colonna
[11].
The novelty of this paper is twofold. First, in Section II
we propose a novel modelling strategy to model sequential
turbocharging with a mean value first principle approach that
does not require detailed compressor and turbine maps, as
an extension to the mean value first principle diesel engine
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S HIP PROPULSION SYSTEM MODELS
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(2)
pitch actuation system is represented with a linear first order
system, based on the analysis in [22], as opposed to more
complex non-linear models proposed in [23]–[25].
(6)
(5)
(7)
MG
G
(3)
Legend:
(3)
(2) Hull
(3) Diesel generator
(3)
(4) Electric drive
(5) Main diesel engine
(6) Gearbox
(7) Shaft
(8) Controllable pitch propeller
(4)
(8)
(4)
(5) In this paper, for the hull model, we consider the
equation of motion in surge direction only, as opposed to
6 degree of freedom models used in Schulten [12] and
Martelli [26], because the propulsion architecture and control
strategy primarily influence acceleration and deceleration in
surge direction. Moreover, we use semi-empirical resistance
estimates performed at the Defence Materiel Organisation
based on Holtrop and Mennen(1982) [27] and improved based
on empirical data from Royal Netherlands Navy frigates.
G
(4)
(4)
G
MG
(8)
(7)
(5)
(6)
Fig. 2: Hybrid propulsion system layout for the case study frigate.
Xset
Tset / nset
Speed Nvirt Control
setpoint
actions
Diesel
engine
Mde
Mp
nde
np
Mem
Electric
drive
Gearbox
and
shaftline
T
Propeller
nem
P
Tset / nset
1) Sequential turbocharger modelling: The turbocharger
performance in this work is based on Zinner blowdown (1),
[13], [28], slip assuming isentropic flow through a nozzle
(4), [13], the elliptic law (5) [13], [29] and assumed balance
between work of the compressor and turbine according to
the Büchi equation (6), [13]. The model description and
validation have been presented in Geertsma et al [9]. The
resulting equations to determine turbocharger performance can
be summarised as follows:
1
(nbld − 1) pd,s (t)
Tbld (t) =
T6 (t)
+
(1)
nbld
nbld
p6 (t)
Td (t) =
(2)
cpg Tbld (t) (m1 (t) + mf (t)) + cpa Tslip ssc (t)m1 (t)
Hull
Electric
drive
Mem
nem
Mde
Xset
(6) The power generation system is represented as a
constant efficiency, with a constant specific fuel consumption
for the diesel generator, because the impact of the power generation system architecture and components is not investigated.
vs
Diesel
engine
Gearbox
and
shaftline
Mp
np
T
Propeller
vs
nde
Fig. 3: Schematic representation of hybrid diesel-electric propulsion
for notional Frigate showing coupling between models.
The hybrid diesel-electric propulsion model consists of 6
sub-models:
cpg (m1 (t) + mf (t)) + cpa ssc (t)m1 (t)
ne m1nom Ψsc (t)
ssc (t) = sscnom nom
(3)
ne (t) m1 (t) Ψscnom
2
κg +1
2κg pd (t) κg
pd (t) κg
−
(4)
Ψsc (t) =
κg − 1
p1 (t)
p1 (t)
2
(ssc (t)ṁ1 (t) + ṁf (t)) Rg Td (t)
+ p2e
(5)
pd,s (t) =
2
Aeff
πcoms (t) =
(6)
⎞⎞
⎛
⎛
1
⎝1 + δf (t)χg ηTC (t)rTTC (t) ⎝1 −
κ −1 ⎠⎠ ,
g
πtur (t) κg
(1) The diesel engine model uses a mean value first
principle (MVFP) approach using the six point Seiliger cycle
proposed in Schulten [12] and evaluates the turbocharger performance assuming balance between compressor and turbine
work according to the Büchi balance [13, Ch. 8]. Therefore,
the model requires significantly less calibration effort than
alternative MVFP diesel engine models that use compressor
and turbine maps [12], [14], [15] or even more complex crank
angle models [16], [17]. The full model description, calibration
and validation have been described in Geertsma et al [9]. In
Section II-A1 the exhaust receiver and turbocharger model
description based on [9] will be revisited and the methodology
to model Sequential Turbocharging (STC) will be introduced.
(2) The induction machine model is a fifth-order state-space
model as proposed in [18], with flux equations and control system modelled in the synchronously rotating reference frame, as
introduced in [10]. The modelling strategy to define the flux
equations and the speed and torque control equations in the
rotating reference frame significantly reduces model run-time.
where Tbld is the Zinner blowdown temperature in K, nbld is
the polytropic expansion coefficient of the blowdown process,
allowing for heat loss, pd,s is the equilibrium pressure in the
exhaust receiver in Pa, p6 and T6 are the pressure in Pa and
temperature in K at point 6 in the Seiliger cycle, when the
exhaust valve opens, Td is the exhaust receiver temperature in
K, cpg and cpa are the specific heats at constant pressure of
the exhaust gas and air in J/kgK, m1 and mf are the mass
of trapped air and fuel per cylinder per cycle in kg, Tslip is
the temperature of the air slip during scavenging allowing for
heat pick-up in K, ssc is the slip ratio of the scavenge process,
relative to trapped air, ne and nenom are the actual and nominal
engine speed in Hz, Ψsc and Ψscnom are the actual and nominal
non-dimensional scavenge flow, p1 is the charge air pressure
(3) The gearbox model predicts the gearbox losses with a
linear torque loss model, based on a detailed thermal analysis
proposed in Godjevac et al [19]. The parameters are determined for the complete gearbox operating envelop as proposed
and validated in Geertsma et al [8], [9].
(4) The propeller model uses the widely accepted four
quadrant open water diagram and, in particular the recently
developed Wageningen C- and D- series for controllable pitch
propellers [20], [21], as described in [8], [9]. Moreover, the
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in Pa, κg is the specific heat ratio of the exhaust gas, pe is the
pressure after the turbocharger in Pa, Aeff is the effective area
of the turbine in m2 , πcom is the static compressor pressure
ratio, δf is the fuel addition factor, χg is the ratio between
the specific heats at constant pressure of the exhaust gas cpg
relative to air cpa , ηTC is the turbocharger efficiency, rTTC is the
driving temperature ratio of the turbocharger, pamb and Tamb are
the ambient pressure in Pa and temperature in K and πtur is
the turbine pressure ratio.
ODact
Fuel
valve
Qta
low
pressure
compressor
Ngt
Qta
Qss
N1
N1
Qe
Power
turbine
Mgt
high
pressure
compressor
pHP7
T6
Fig. 4: Schematic representation of gasturbine simulation model that
replaces the diesel engine in Figure 3 for hybrid gas-turbine-electric
propulsion analysis.
The novelty of the model in this paper is that Sequential
Turbocharging (STC) with multiple turbochargers is represented by changing the effective turbocharger area Aeff in
accordance with the number and size of the running turbochargers relative to the turbochargers running at nominal
speed. In the case study for example, reducing the effective
turbocharger area Aeff with a factor 2 below 750 rpm represents
the engine switching from two to one turbochargers below 750
rpm. As proposed in [9] the nominal effective turbocharger
area Aeff is determined with nominal performance value and
(5). Moreover, the non-dimensional scavenge flow is increased
proportional to the air flow increase during opening of the
cylinder bypass valve.
TABLE I: Frigate case study diesel engine model parameters.
nominal engine power Penom
nominal engine speed nenom
number of cylinders ie
number of revolutions per cycle ke
bore diameter DB
stroke length LS
crank rod length LCR
crank angle after TDC, inlet closure αIC
crank angle after TDC, exh open αEO
nominal spec. fuel cons. mbsfcnom
heat release efficiency ηq
geometric compression ratio εc
total nominal mass flow ṁtnom
cylinder volume at state 1 V1
nominal pressure at state 1 p1nom
maximum cylinder pressure pmaxnom
temperature after the intercooler Tc
temperature of the inlet duct Tinl
parasitic heat exch effectiveness εinl
fuel injection time delay τX
turbocharger time constant τTC
exhaust receiver time constant τpd
gas constant of air Ra
specific heat at constant vol of air cv−a
specific heat at const. press of air cpa
specific heat at const. p of exhaust cpg
isentropic index of air κa
isentropic index of the exhaust gas κg
lower heating value of fuel hL
stoichiometric air to fuel ratio σf
polytropic exponent for expansion nexp
polytropic exponent for blowdown nbld
nominal mechanical efficiency ηmnom
constant volume portion grad Xcvgrad
constant temperature portion Xctnom
turbocharger factor aη
turbocharger factor bη
turbocharger factor cη
ambient pressure pamb
ambient temperature Tamb
B. Hybrid gas-turbine-electric propulsion
The hybrid gas-turbine-electric model uses an empirical
model developed to perform studies on the ship control system
and dynamic ship performance of the dutch air-defence and
command frigates (LCF) [30]. This model uses data of the
Rolls Royce Spey SM1C, a marinised version of the Rolls
Royce Spey aircraft engine. A dynamic model is not required
for the analysis of the gas-turbine as we are not considering
thermal loading of the gas-turbine, Alternatively, we use the
acceleration rate that has been prescribed by the manufacturer.
The simulation model of the Spey SM1C consists of 4
sub-models, see Figure 4. The fuel valve submodel defines
the fuel flow QTA based on the actual position of the fuel
valve ODact . The LP-compressor submodel predicts the speed
of the LP-compressor N1 based on the fuel flow QTA . It also
predicts the energy needed to drive the gas generator QSS
and the remaining energy leaving the gas generator in the
exhaust gases Qe . The power turbine submodel determines
torque delivered by the power turbine Mgt as a function of
fuel flow and energy required by the gas generator as well as
rotational speed of the power turbine Ngt and LP-compressor
N1 . The HP-compressor submodel, finally, predicts the inlet
temperature of the power turbine T6 and the HP-pressure pHP7
as a function of LP-compressor speed N1 and the energy of
the exhaust gases Qe .
C. Model parameters
The model parameters used for the hybrid diesel-electric
propulsion model are presented in Table I for the diesel engine,
in Table II for the induction machine, in Table III for the
gearbox, in Table IV for the propeller and in Table V for the
hull. Moreover, the estimated resistance of the hull for trial,
design and off-design conditions is presented in Figure 5.
9100 kW
16.7 rev/s
20
2
0.28 m
0.33 m
0.64063
224 °
119 °
189 g/kWh
0.915
13.8
17.26 kg/s
0.0199 m3
4.52e5 Pa
206e5 Pa
323 K
423K
0.05
0.015 s
5s
0.01 s
287 J/kgK
717.5 J/kgK
1005 J/kgK
1100 J/kgK
1.4
1.353
42700 J/kg
14.5
1.38
1.38
0.90
-0.4560
0.4
-5.13e-12
-3.99e-6
0.092
1e5 Pa
318 K
below the operating speed at which the second turbocharger
is switched of, 750rpm. Similarly the engine air acces ratio
steeply increases below 750 rpm, leading to a higher power
limit in part load, because the increased charge pressure in
The resulting diesel engine fuel consumption and air access
ratio over the full operating range are illustrated in Figure
6. The specific fuel consumption clearly drops significantly
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TABLE II: Induction machine parameters
trial condition
service condition
off-design condition
5
3150 V
66.5 rad/s
10.30 Ω
0.534 Ω
0.2522 Ω
0.0630 Ω
0.0552 Ω
3000 kW
2000
ship resistance [kN]
0.0269
0.7254
0.2454
484 (4%)
7.752
129
0
0
5
15
20
25
30
35
Fig. 5: Predicted ship resistance in trial, service and off-design
condition.
brake specific fuel consumption contour [g/kWh]
9000
800
speed [rpm]
12.9
2.2
4
2.2
300
400
600
1000
2 1.9
2.
2.4 2
2.28.6
1000
600
2.2
3
19
0
119
29089
0
2000
240
260
0
400
2.4
2.6
2.8
4 33.5
205
220
0
400
2.8
3000
205 210
2
192
67
195 1919
3.5
power [kW]
4000
3
188
5000
210
220
240
260
300
400
600
1000
5200
2
0.155
208
1998
5 11
19
6
18
2.6
4000
6000
2.4
power [kW]
5000
2.8
7000
6
18
6000
2000
2.62.
8000
7000
TABLE V: Hull model parameters.
4
power limit [kW]
air excess ratio [-]
11119
11811929
21098995067890
86
8000
2.6
2.8
33.5
0.09
1
4.8
1.4
1.8
0.238
1.67
1
0.0524
air excess ratio contour plot
9000
power limit [kW]
bsfc [g/kWh]
3000
ship mass m in 103 kg
number of propellers m
thrust deduction factor t
10
ship speed [kts]
TABLE IV: Propeller parameters
wake fraction w
relative rotative efficiency ηR
propeller diameter D in m
design pitch ratio at 0.7R Pd
nominal pitch ratio at 0.7R Pnom
pitch ratio for zero thrust P0
first order pitch actuation delay τP
Vrijdag coefficient c1
shock free entry angle αi
1000
500
TABLE III: Gearbox parameters
gearbox loss parameter agb
gearbox loss parameter bgb
gearbox loss parameter cgb
gearbox nominal power loss Plnom in kW
gearbox speed reduction ratio igb
nominal propeller speed npnom in rpm
1500
20
019
919
819
71
96
pole pairs P
nominal voltage V
base speed ωb
mutual reactance xm
stator self reactance xs
rotor self reactance xr
stator resistance rs
rotor resistance rr
nominal power Pnom
resistance from model tests corrected for envronmental conditions and fouling
2500
3 3.5
4
5
600
800
1000
speed [rpm]
Fig. 6: Diesel engine model specific fuel consumption and air excess
ratio results in complete operating envelope.
part load from the single turbocharger increases the air access
ratio and therefore reduces thermal loading. Due to a lack of
Factory Acceptance test data on an engine with Sequential
Turbocharging, these results have not been validated. The
resulting specific fuel consumption over the entire operating
profile for the gasturbine is illustrated in Figure 7.
III.
electric propulsion and hybrid diesel-electric propulsion.
The hybrid gas-turbine-electric propulsion system will
consist of two modes: electric mode and gas-turbine mode.
The hybrid gas-turbine-electric propulsion system will run in
electric mode up to 18 kts, the maximum speed that can be
attained on the electric drive of 3MW with a control objective
to run silently and to run with low fuel consumption and
emissions. When the ship needs to sail at speeds in excess of
18 kts, the system will run in gas-turbine mode. The control
objective for this mode is high manoeuvrability and top speed.
Because the gas-turbines have a significantly higher specific
fuel consumption and the electric drive can run at rated torque
at all speeds, running the gas-turbine and electric drive in
C ONTROL STRATEGIES
Depending on the ships task or mission, different control
objectives can be required. For example, during a transit the
ships fuel consumption and emissions should be as low as
possible to extend the ships operational range and reduce
impact on the environment, while during high air threat the
ship should be highly manoeuvrable to position itself and
optimally deploy radar and weapon systems. Therefore, we
consider different control modes for both hybrid gas-turbine-
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Specific Fuel Consumption RR Spey SM1C
shaft speed [%], relative torque [%] and relative pitch setpoint [%]
250
260
14000
300
280
12000
350
250
260
400
10000
280
450
8000
6000
300
500
Engine Power [kW]
combinator curves for manoevre, transit and electric modes
100
1
23
240
16000
235
propeller curve
upper engine envelope
lower engine envelope
specific fuel consumption
233
18000
300
350
4000
350
400
2000
0
500
600
1000
1500
400
450
500
450
500
600
2000
2500
3000
3500
600
700
4000
4500
5000
5500
6000
80
60
40
20
0
-20
-40
-80
-100
-60
Power Turbine Speed [rpm]
shaft speed for manoeuver mode
relative pitch setpoint for manoevre mode
relative torque setpoint for manoeuvre mode
shaft speed for transit mode
relative pitch setpoint for transit mode
electric drive speed setpoint for electric mode
-60
-40
-20
0
20
40
60
80
100
120
virtual shaft speed setpoint [rpm]
Fig. 7: Specific fuel consumption from empirical gas-turbine model.
Fig. 8: Combinator curves for manoeuvre and transit mode of hybrid
diesel-electric propulsion and electric mode of hybrid gas-turbine
electric propulsion.
parallel has hardly any advantages, apart from the possibility
to increase acceleration performance on gas-turbines, which is
already considered sufficient.
TABLE VI: Control parameters in manoeuvre (Man), fast manoeu-
The hybrid diesel-electric propulsion system will run the
diesel engine and electric drive in parallel. In transit mode the
control objective is to minimise fuel consumption and limit
engine thermal loading to reduce required maintenance. In
manoeuvre mode, the control objective is to accelerate the ship
as fast as possibly for optimum manoeuvrability, while not
thermally overloading the engine and prevent engine failure
and excessive maintenance.
vre (Man2), transit (Tran) and electric (Elec) mode
mode
proportional gain speed KPS
reset rate speed KIS
proportional gain torque KPT
reset rate torque KIT
proportional gain field KPD
reset rate field KID
rel. pitch incr. rate (%/s2 )
rel. torque incr. rate (%/s2 )
rel. speed incr. rate (%/s2 )
A. Gas-turbine propulsion
The engine controller submodel controls the angle of opening of the fuel valve, referred to as throttle demand ODdemand .
The position of the fuel valve depends on inlet temperature of
the power turbine T6 , the rotational speed of the power turbine
Ngt and low pressure (LP)-compressor N1 and the demanded
LP-compressor speed N1demand . During the acceleration of the
gas generator, the fuel flow increase rate rQta is limited to
prevent stall. The engine controller requires the pressure after
the high-pressure (HP) compressor HP7 to control this rate.
In the case study, the limiting factor for acceleration are fuel
flow increase rate rQta , at 1%/s, and relative pitch increase rate
of 0.56%/s. This acceleration rate has been confirmed during
recent measurements on board LCF De Zeven Provinciën.
Man
2
0.5
10
1
160
5
0.56
1.11
0.6
Man2
10
1
10
0
160
5
2.5
1.25
10
Tran
2
0.5
10
1
160
5
0.21
0.42
0.2
Elec
10
1
10
0
160
5
5.56
11.1
5.6
actuation system. The increase rate of the electric drive limits
dynamic loading of the power generation system. The PID
parameters have been determined by trial and error, using
Ziegler-Nichols [31]. The control parameters are listed in Table
VI and the combinator curve is illustrated in Figure 8.
In this study, the control strategy was modelled in the
synchronously rotating reference frame, like the induction
machine in [10]. While in a real system the actual quadrature
and direct current in the synchronously rotating reference
frame ieqs and ieds can be determined from flux and stator
current measurements as discussed in [10], we directly use
the quadrature and direct current in the synchronously rotating
reference frame from the simulation model. Thus, simulation
time is significantly reduced, due to the absence of fluctuating
sine wave signals. This can be allowed, because we are
not interested in the effects of noise and inaccuracy of the
measurements. The resulting PID torque control equations are
described in [10].
B. Electric propulsion
The control strategy in electric mode aims to provide fuelefficient, silent propulsion and consists of a combinator curve
that determines shaft speed and propeller pitch setpoints from
the requested virtual shaft speed, PID control on shaft speed for
the electric drive, and feedforward control for propeller pitch.
Within the speed control loop, the electric drive uses field
oriented torque control, as discussed in Geertsma et al [10]
and Ong [18]. The increase rate of propeller pitch is limited
to reflect the maximum pitch change rate of the hydraulic
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Nvirt
Combinator curve
Tset
PIcontrol
PIcontrol
Xset
nde
vs
e
Fuel consumption as a function of ship speed
250
Pset
limits
is
e
200
Plim
fuel consumption [kg/mile]
nset
Propulsion model – Figure 3
Fig. 9: Schematic representation of manoeuvre and transit control
strategies for hybrid diesel-electric propulsion.
150
100
50
C. Hybrid diesel-electric propulsion in manoeuvre mode
manoeuvre mode with electric assist
transit mode with power take-off
electric drive mode
gas-turbine mode
In manoeuvre mode, the control objective is to minimise
acceleration time without thermally overloading the engine.
In order to achieve this, the electric drive is used to support
the diesel engine. The first control strategy consists of a
combinator curve that determines shaft speed, electric drive
torque and propeller pitch setpoints from the requested virtual
shaft speed and PID-control on electric drive torque and engine
speed, and feedforward control on propeller pitch, as illustrated
in Figure 9. The increase rate of propeller pitch is limited to
reflect maximum pitch change rate of the hydraulic actuation
system. The increase rate of the diesel engine speed limits
thermal loading in accordance with the engine project guide
[32]. The electric drive torque control uses field oriented
control as detailed in [10], directly using the quadrature and
direct current in the synchronously rotating reference frame
ieqs and ieds from the simulation model as discussed in Section
III-B. Engine speed control and propeller pitch control are
described in more detail in [9]. In the second control strategy,
fast manoeuvre mode, the electric drive runs at speed control
and the diesel engine runs at torque control as proposed in
[10], [33], with the speed setpoint for the electric drive and
the torque setpoint for the diesel engine defined in the same
combinator curve. Because the electric drive is not sensitive
for fast torque increases and the diesel generator can accelerate
more quickly because it runs at constant speed, the torque
increase rate can be higher. Moreover, because the electric
drive accelerates shaft speed faster, engine charge air pressure
increases faster, and therefore the diesel engine can be loaded
faster at lower thermal loading. The control parameters are
listed in Table VI and the combinator curve is illustrated in
Figure 8.
0
0
5
10
15
20
25
30
ship speed [kts]
Fig. 10: Fuel consumption as a function of ship speed in manoeuvre,
transit, electric and gas-turbine mode of hybrid diesel-electric and
gas-turbine-electric propulsion.
IV.
F RIGATE CASE STUDY RESULTS
The performance metrics proposed in [9] will be used for
performance evaluation: (1) Fuel consumption per mile for
trial, design and off-design conditions, presented as a function
of ships speed in Figure 10; (2) Average air excess ratio
at constant speed for trial, design and off-design conditions,
presented as a function of ship speed, which serves as an
indicator for engine thermal loading during constant speed
sailing due to the average temperature in Figure 11; (2)
Acceleration time for speed increases from 0 to 10, 10 to
20 and 20 to 25, when increasing virtual shaft speed to the
required speed, and from 0 to 25 kts, when increasing virtual
shaft speed to the maximum setting, in design conditions in
Table VII; (4) Minimum air acces ratio during speed increases
from 0 to 10, 10 to 20 and 20 to 25 with a step increase
and from 0 to 25 kts, when increasing virtual shaft speed to
the maximum setting, in design conditions in Table VIII. This
performance criterion serves as an indicator for engine thermal
loading due to acceleration manoeuvres; (5) Cavitation plot
of acceleration manoeuvres from 0 to 10 and 10 to 15 kts
in design conditions in Figure 12 as proposed in [34]. The
MATLAB Simulink R2106b software has been used on a PC
with Intel Core i7 processor and 16 GB memory to simulate the
propulsion plants. A typical 9000 s simulation of 4 acceleration
manoeuvres takes around 7 s.
D. Hybrid diesel-electric propulsion in transit mode
In transit mode, the control objective is to minimise fuel
consumption and emissions. In order to achieve this, the
electric drive is used as a generator, as power take-off, because
the specific fuel consumption (SFC) of the main engine is
lower than that of the generator, and the main diesel engine
with sequential turbocharging has an operating envelope that
enables increased power in part load. Moreover, the power
increase due to power take-off reduces the main engine SFC
even further, as illustrated in Figure 6. Apart from a combinator
curve with power take-off and other increase rate settings, the
control strategy is the same as described in Section III-C. The
control parameters are listed in Table VI and the combinator
curve is illustrated in Figure 8.
A. Discussion
First, hybrid gas-turbine electric propulsion achieves 3 kts
higher top speed than hybrid diesel-electric propulsion. The
greater engine power of the gas-turbine driven ship causes this.
Second, the acceleration time of the diesel-electric propulsion
system in manoeuvre mode is 208% longer than that of gasturbine-electric propulsion. This has two reasons: first, the
gas-turbine delivers 1.5 times the power of the diesel engine
with electric drive and, second, the diesel engine accelerates
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Air access ratio as a function of ship speed
Cavitation plot
14
4
12
3.5
10
Cavitation number
air excess ratio
4.5
3
2.5
8
6
2
notional cavitation curve
reduced pitch combinator transit 0 to 10 kts
reduced pitch combinator transit 10 to 15 kts
manoeuvre mode 0 to 10 kts
manoevre mode 10 to 15 kts
transit mode 0 to 10 kts
transit mode 10 to 15 kts
4
manoeuvre mode with electric assist
transit mode with power take-off
1.5
2
0
5
10
15
20
25
30
2
4
ship speed [kts]
6
8
10
12
14
effective angle of attack
Fig. 11: Air excess ratio λ as a function of ships speed in manoeuvre
and transit of hybrid diesel-electric propulsion.
Fig. 12: Cavitation plot as defined in Geertsma [9] for acceleration
manoeuvres in transit mode, manoeuvre mode and transit mode with
reduced pitch combinator curve.
TABLE VII: Acceleration time and top speed for acceleration in
manoeuvring (Man) fast manoeuvring (Man2), transit (Trans), Electric
(Elec) and Gas-turbine (Gas-t) mode.
speeds (kts)
From 0 to 25
From 0 to 10
From 10 to 20
From 20 to 25
Top speed
Man
181 s
261 s
150 s
87 s
26
Man2
104
187
98
58
26
Trans
448 s
463 s
297 s
166 s
26
Elec
177 s
18
of 18 kts the gas-turbine-electric system uses 10% more fuel
in electric mode. The use of power-take-off accounts for 7%
of this fuel reduction. At speeds below 14 kts electric mode
is more efficient, because the electric drive can run at a
lower shaft speed, while the diesel engine is limited to its
minimum speed of 400 rpm. Thus, the pitch needs to be
reduced when the engine is driving the shaft and open water
efficiency significantly drops. Moreover, the thermal loading of
the engine, in this study quantified with the air excess ratio λ,
stays within acceptable limits, above 1.85, when using powertake-off in the very efficient transit mode. Finally, the large
pitch ratio of the propeller leads to a high angle of attack,
potentially increasing risk of cavitation occurring. This risk
of cavitation can be reduced by reducing the propeller pitch
in silent mode and reducing the increase rate of propeller
pitch and engine speed. the reduced propeller pitch leads
to a reduced angle of attack, reduced propeller loading and
increased propeller speed, as illustrated with the reduced pitch
combinator curve with 70 % pitch in Figure 12.
Gas-t
87
339
131
74
29
TABLE VIII: Minimum air excess ratio λ for acceleration in
manoeuvring (Man), fast Manoeuvring (Man2) and transit (Trans)
mode (design condition).
speeds (kts)
From 0 to 25
From 0 to 10
From 10 to 20
From 20 to 25
Man
1.85
2.35
2.24
2.18
Man2
2.35
2.88
2.20
2.33
Trans
2.04
1.90
1.93
2.17
V.
C ONCLUSION AND FUTURE RESEARCH
This paper has investigated the trade-off between reducing
fuel consumption and emissions of ships and achieving good
acceleration performance and ship top speed at an acceptable
thermal loading of the engine. The work has illustrated that the
control strategy has a very significant contribution to the ships
performance. The case study has demonstrated that, while
top speed and acceleration performance of gas-turbine driven
ships are better, parallel hybrid diesel-electric propulsion can
reduce fuel consumption with 25% at top speed and 10% at
transit speed. It can also provide an acceleration performance
that is significantly better than diesel-mechanical propulsion.
However, hybrid diesel-electric speed control with the electric
drive running at speed control and the diesel engine at torque
control can potentially achieve acceleration performance that
nearly measures up to gas-turbine acceleration. Moreover, in
in 180 s to 100% power while the gas-turbine rates up to
90% in 90 s. The diesel engine can however be accelerated
faster without increasing thermal loading if the electric drive
runs at speed control and the engine at torque control. Then
acceleration to full speed takes only 20% longer. Moreover,
the propulsion plant can use either more high-speed engines or
larger medium-speed engines to meet the propulsion power of
gasturbines at the cost of extra weight. Finally, adaptive pitch
control strategies can further improve acceleration performance
of hybrid diesel-electric propulsion [7].
On the other hand, at a speed of 26 kts the gas-turbineelectric propulsion system uses 25% more fuel and produces
equally more fuel-related emissions, and at a transit speed
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order to achieve acceptable cavitation performance with an
efficient high pitch propeller, controllable pitch propellers and
adaptive pitch control can reduce the angle of attack in silent
mode, reducing the risk of cavitation.
[11]
[12]
In this work, engine thermal loading is quantified with
the air excess ratio. More work is required to establish the
thermal loading limits and to determine how fast diesel engines
can accelerate with torque control and adaptive pitch control strategies, without causing increased engine maintenance.
While thermal loading of the generators was not considered in
this study, electric drives can potentially cause excessive load
fluctuation on the generators. Future studies should investigate
whether batteries can provide a means of load levelling or peak
shaving and further improve fuel efficiency by running diesel
generators in more efficient operating points. Finally, this work
demonstrated the effect of various combinator curves. More
research is required to determine optimum acceleration and
cavitation behaviour with adaptive pitch control strategies.
[13]
[14]
[15]
[16]
[17]
In conclusion, this work has shown that a 20% reduction of
fossil fuel dependency compared to current hybrid gas-turbineelectric propulsion is feasible almost completely by applying
hybrid diesel-electric propulsion and using power take-off.
While this will cause a slight reduction in top speed, advanced
control strategies can ensure acceleration performance similar
to gas-turbine propulsion.
[18]
[19]
[20]
ACKNOWLEDGMENT
This research is supported by the project ‘ShipDrive: A Novel
Methodology for Integrated Modelling, Control, and Optimization of
Hybrid Ship Systems’ (project 13276) of the Netherlands Organisation for Scientific Research (NWO), domain Applied and Engineering
Sciences (TTW). The Royal Netherlands Navy supplied Figure 1.
[21]
[22]
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