Paper ID: ETC2019-004 Proceedings of 13th European Conference on Turbomachinery Fluid dynamics & Thermodynamics ETC13, April 8-12, 2018; Lausanne, Switzerland

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PERFORMANCE IMPROVEMENT OF THE CFM56-3 AIRCRAFT

ENGINE BY ELECTRIC POWER TRANSFER

H. Balaghi Enalou, S. Bozhko

University of Nottingham, Nottingham, the UK

hbalaghienalou@gmail.com, Serhiy.bozhko@nottingham.ac.uk

ABSTRACT

With the design trends towards the More Electric Engine (MEE) for the More Electric

Aircraft (MEA), areas for novel technologies can be pinpointed for multi-spool engines,

introducing remarkable improvements to push the boundaries of propulsion technology as it

strives to create quieter and more efficient engine. Provided that a multi-spool engine is

equipped with electrical machines connected to each of its shafts, using power electronics within

a single high-voltage DC bus configuration, it is possible to circulate desired amount of power

between the engine shafts independent of their speeds. This paper presents an engine model

which has also considered Variable Stator Vanes (VSVs) and Variable Bleed Valves (VBVs) for

bleeding in order to investigate the idea of power transfer at low speed settings. Validation with

test results from the CFM56-3 engine highlights acceptable level of accuracy of the engine

model. Moreover, preliminary results show that the idea of power circulation is highly desirable

for low speed settings for the CFM56-3 aircraft engine. The electrical power transfer from the

Low Pressure (LP) to the High Pressure (HP) shaft at engine’s low speed settings such as taxi

and flight idle, helps to decrease the fuel rate and increase the available surge margin of

compressors.

KEYWORDS

AIRCRAFT ENGINE, ELECTRIC POWER TRANSFER, MORE ELECTRIC ENGINE

(MEE), VARIABLE BLEED VALVES (VBVS), VARIABLE STATOR VANES (VSVS)

NOMENCLATURE

CC Combustion Chamber PEC Power electronic Convertor

CV Control Volume SFC Specific Fuel Consumption (g.kN-1.s-1)

ECS Environmental Control System SLS Sea-Level-Static

EEC Electronic Engine Control SM Surge Margin

EOM Engine Operation Mode VBVs Variable Bleed Valves

EPS Electrical Power System VSVs Variable Stator Vanes

EPT Electric Power Transfer 𝐽 Shaft inertia (kg.m2)

HP High Pressure 𝑃 Power (Watt)

HPC High Pressure Compressor 𝑅 specific gas constant (J.kg-1.K-1) HPT High Pressure Turbine 𝑇 Temperature (°K) ICV Intermediate Control Volume 𝑉 Volume (m3) LP Low Pressure

𝑚 Mass (kg)

LPC Low Pressure Compressor 𝑝 Pressure (pa) LPT Low Pressure Turbine 𝑡 Time (s)

MEA More Electric Aircraft 𝜔 Speed (rad.s-1)

MEE More Electric Engine

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INTRODUCTION

Conventional multi-spool gas turbines can operate within a wide speed range for industrial,

propulsion or power generation purposes. Among them, high bypass ratio turbofans are considered

as the dominant sources of propulsion for most of civil aircrafts. Recently, the MEE has encouraged

the use of high-power, high-efficiency, lightweight motor/generators connected to both the Low

Pressure (LP) and High Pressure (HP) of the engine to augment electric power to the airframe of the

MEA and reduce the effects of large load transients (Provost, 2002, Pluijms et al., 2008, Kloos et al.,

2018). Having parallel sources of power with alternating speed on various engine shafts has resulted

in introduction of innovative power system architectures with Power Electronic Convertors (PECs)

controlling the generation. Engine designers have a great opportunity to take advantage of these

architectures, introducing new integrated systems within the MEE concept.

However, the off-design efficiency of multi-spool engines such as aircraft engines is very low,

irrespective of the engine size, in part because of the particular manner in which the engine operates

at these conditions. Typically, for gas turbines, higher engine pressure ratio leads to higher efficiency,

while at part loads, the engine pressure ratio drops due to the decrease in the speed of compressors.

The engine components (compressors and turbines) are thermodynamically coupled since the Low

Pressure Compressor (LPC) delivers airflow to the High Pressure Compressor (HPC) and the High

Pressure Turbine (HPT) delivers the hot combustion products to the Low Pressure Turbine (LPT).

This makes an approximately fixed power split ratio between the HPT and the LPT which, as a result,

forces the shaft speed variations to bind to each other thermodynamically from part to full power

settings, although there is no mechanical link between the shafts. The inflexibility of shaft speeds

causes mismatches in the performance of compressors at off-design conditions which is dealt with by

handling bleed systems such as Variable Bleed Valves (VBVs) between the compressors. As a

consequence, the engine operates in a sub-optimal condition at low compressor speeds, thereby being

fuel inefficient.

In this case, any system which can help to decouple the shaft speeds is highly desirable. Power

transfer has been highly desirable since 1990 for engineers, suggesting various methods such as

implementation of a hydraulic transmission system with hydraulic pumps and motors, multi speed

transmission systems with clutch assemblies, auxiliary air turbines, electric heaters and magnetic

gearboxes (Hield et al., 1993, Eick et al., 2005, Anghel et al., 2012). Among all, provided that there

is at least one electrical machine on each shaft, there is the ability to transfer power between the shafts

with Power Electronic Convertors (PECs) with high efficiency and reliability.

This idea can be applied broadly for applications which include, but are not limited to, aircraft

engines, ship engines and ground-based power generations. The main expectations from this

configuration are:

• fuel burn reduction at off-optimum conditions

• increase in engine stability margins

• dynamic response improvement

Moreover, this configuration will decouple shaft speeds to some degree, depending on the amount

of power transfer to shaft powers, in order to:

• improve engine starting and relighting

• optimize compressors’ performance at higher efficiency with adequate Surge Margin (SM)

• remove Variable Stator Vanes (VSVs) and Variable Bleed Valves (VBVs)

• help faster acceleration and deceleration of the engine;

• help the engine core withstand severe transients

• increase the engine’s thermodynamic cycle efficiency

• manage engine start-up and performance at very high or low ambient temperature

Prior control schemes, mainly suggested for turbofans, have elaborated on multi-spool offtake for

the More Electric Aircraft (MEA), attempting to optimize engine performance by splitting the

electrical loads on the two spools (Pluijms et al., 2008) or circulation of power between them to

further optimize and enhance the engine performance (Enalou et al., 2018, Phillips et al., 2013).

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(Belokon et al., 2003) has also considered power exchange with a LP shaft through an auxiliary

generator/motor to maintain a high peak temperature to optimize engine efficiency at part-load

conditions in a multi-spool turbo-generator for a distributed power generation system.

This paper investigates fuel burn and SM of the engine by implementing the concept of power

exchange for the CFM56-3 engine model which has considered the effect of VBVs and VSVs. For

this purpose, a brief overview on multi-spool engine performance will be provided. Then the electric

power transfer system will be introduced. Afterwards, the engine model will be elaborated followed

by performance results at different operation scenarios.

MULTI-SPOOL ENGINE

Figure 1 shows the schematic diagram of a typical high bypass turbofan where the LP spool is

colored as blue and the HP spool is colored as dark brown. The common components of these engines

can be named as: fan, LPC, inter-compressor bleeds, HPC and cooling bleeds, fuel metering system,

combustion chamber, HPT, LPT and discharge nozzles. The HPT drives the HPC on the HP shaft and

the LPT drives the fan and LPC on the LP shaft.

For the multi-spool engine,

corrected speed of the HPC

determines the outlet conditions of the

LPC. The HPC will swallow and pass

flow delivered by the LPC, provided

that the HPC speed is properly

matched with that of the LPC.

However, if the HPC speed is less than

its desired matched value with that of

the LPC, the HPC demands less mass

flow. Hence, it acts as a blockage for

the rear side of the LPC, pushing it to

decrease its mass flow on the same

speed line as shown in Figure 2. As a

result, the pressure ratio of the LPC

increases and the operation point of

the LPC gets closer to the surge line.

Therefore, the operating line of the

LPC is determined by the flow

requirement of the HPC which is

determined by the HPC corrected

speed. LPC and HPC speeds are

normally matched for the high speed

setting at design condition while, due

to the HPT/LPT power split ratio at

lower speed settings, the HPC speed is

lower than the desired matched value.

For this reason, there is a risk of LPC

surge at low speed settings.

To prevent this, the VBVs are

located between the LPC and HPC

with the function of regulating the

primary flow entering the HPC. VBVs

act primarily at low speeds due to

lower LPC SM. However, bleeding of

FAN

LPC

HPCCC

HPT

LPT

Core Nozzle

Duct Nozzle

Intake

Fuel

Flow direction

GLP

SGHP

ACDC

ACDC

DC bus

MEADistribution system

Figure 1 Schematic diagram of a turbofan

Corrected massflow Figure 2. Operational speed line on LPC map

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compressed air is a loss of energy, for which VBVs positions are normally scheduled with the HPC

speed to avoid considerable drop in efficiency. The HPC operation point also moves closer to the

surge line with reduction in its speed due to considerable power offtake (Enalou et al., 2016). By

power offtake from the LP shaft and feeding it to the HP shaft the speed reduction will be

compensated. Moreover, with higher HP speeds HPC can afford more massflow which is accordingly

provided by scheduled VBVs. The decrease in bleeding will also increase the engine efficiency. This

highlights the fact that power transfer to HP shaft in order to increase its speed is highly desirable.

The Variable Stator Vanes (VSVs) are also located in the first three stages of the HPC. They form

the stators of these stages. Their objective is to optimize the HPC performance for each combined

state of HP spool speed and air density.

ELECTRIC POWER TRANSFER BETWEEN SHAFTS

Already offered in the MEE (Provost, 2002, Newman, 2004, Hirst et al., 2011, Kreuzer and

Niehuis, 2017), electrical machines are connected to both LP and HP shafts through fixed speed ratio

gearboxes shown in Figure 1. (Gao et al., 2015, Gao and Bozhko, 2016, Gao et al., 2016) have also

elaborated on the single DC bus architecture to handle variable frequency power generation. Figure

3 shows a single DC bus with two electrical machines connected to it through bidirectional converters.

This provides the opportunity to circulate power

between the shafts, having the privilege of

dynamically controlling it. While one of the

electrical machines is controlling the DC link

voltage, the other one can be in current control mode

to manage the amount of power which is circulated

between the engine shafts. This configuration

accommodates an engine with flexible power split

ratios at different operational speeds depending on

the amount of power transfer as described:

𝑃𝐻𝑃𝑇 + 𝑃𝐸𝐻𝑃 = 𝐽𝐻𝑃𝜔𝐻𝑃�̇�𝐻𝑃 + 𝑃𝐻𝑃𝐶 (1)

𝑃𝐿𝑃𝑇 − 𝑃𝐸𝐿𝑃 = 𝐽𝐿𝑃𝜔𝐿𝑃�̇�𝐿𝑃 + 𝑃𝐹𝐴𝑁 + 𝑃𝐿𝑃𝐶 (2)

𝑃𝐸𝐿𝑃 + 𝑃𝐸𝐻𝑃 + 𝑃𝐸𝑃𝑆 = 0 (3)

where; 𝑃𝐸𝐿𝑃 and 𝑃𝐸𝐻𝑃 are electric power transfer for LP and HP shafts, 𝑃𝐸𝑃𝑆 is the onboard MEA

electrical power system load on the DC link. It should be noted that with increasing power demand for

the MEA, the impact of MEA load on the DC link and electric power transfer should be analyzed

together.

MODELLING

A zero-dimensional (0-D) multiple spool engine model has been developed in the MATLAB-

SIMULINK environment by using the ICV method (Enalou et al., 2016) which has been verified in

(Enalou et al., 2017).

In the ICV method, the model includes component, Control Volume (CV), inertia and combustor

modules. It is important to maintain mass/momentum/energy conservation through these modules.

Within each component module, the inputs are inlet and back pressure, inlet temperature and shaft

speed (shown in Figure 4). This data is then used to find the performance point of each respective

component. Components including fan, compressors and turbines are represented by maps in which

corrected mass flow and adiabatic efficiency are indexed by corrected shaft speed and pressure ratio.

Outlet temperature and power can be obtained by having inlet mass flow and efficiency and

implementing thermodynamic laws for open systems including conservation of mass and energy.

Once mass flow and efficiency of every component has been interpolated using component maps, the

outlet temperature is calculated.

GEN1

dc Bus

AC/DC

GEN2

AC/DC

cable

cable

G1

G2

Figure 3 Configuration of single DC

bus multi-source EPS

(Gao et al., 2015)

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Control Volume Component Control Volume

From next component

Power

To shaft inertia module

inT inT

inP

inT

outP

From shaft inertia module

N

inm.

outm. outm

.

inm.

Figure 4 Structure of gas turbine model in ICV method

CV modules consider the flow transient between each two consecutive components. The change

in mass inside the CV is calculated using:

𝑑𝑚𝐶𝑉

𝑑𝑡= �̇�𝑖𝑛 − �̇�𝑜𝑢𝑡

(3)

where; 𝑚𝐶𝑉 is the current mass of the CV. Pressure, Temperature and volume of the mass inside

the CV are related by using the ideal gas law as below:

𝑝𝑉 = 𝑚𝐶𝑉𝑅𝑇 (4)

Pressures and temperatures are total amounts including the kinetic energy of the gas. Assuming

no temperature gradient in the control volume (Rahman and Whidborne, 2009), by using (3) and (4),

the outlet pressure can be determined as:

𝑑𝑝

𝑑𝑡=�̇�𝑖𝑛 − �̇�𝑜𝑢𝑡

𝑉𝑅𝑇⁄

(5)

The imbalanced torque by components on engine shafts leads to acceleration or deceleration

which after integrating over time gives up the shaft speeds. The shaft inertia modules in Simulink,

shown in Figure 5, use Newton’s second law, where; 𝑃𝐺𝑇 is power produced by the turbine and 𝑃𝐿𝑜𝑎𝑑

is power consumed by compressors and other loads on the shaft and J is the shaft inertia.

For the studied engine in this paper, VBVs and

VSVs have been scheduled against HP speed as it is

shown in Figure 6. The deviations of the VSVs from

nominal schedule have a secondary relevance. The

effect of closing the VSVs relative to the nominal

schedule will shift the whole compressor map as

shown in Figure 7. Corrected massflow, pressure ratio of the HPC map are modified in the model to

make it match with the operation line of correlated test results on the HPC map. Efficiency values are

also modified in order to make the simulation results match with the specific fuel consumption of the

correlated test results.

-+

1Js

PLoad

PGT

Figure 5 Inertia module

6

Figure 6 VBVs’ and VSVs’ schedules

(CFM, 1992)

Figure 7 VSVs deviations from nominal

schedule (Kurzke, 2008)

MODEL VALIDATION

The engine model has already been validated for a multi-spool turbo-generator in (Enalou et al.,

2017). This section presents the performance validation for the CFM56-3 engine for which generic

component maps from (GASTURB), that are selected from similar turbomachines and with similar

design, have been implemented for the Off-Design simulation. They are all from axial flow machines,

therefore they are scaled in such a way that they fit to the cycle design point of the CFM56-3 engine.

In order to validate the simulation results of the engine model, compressor operating lines are

compared with the ones from the correlation test report data for the CFM56-3 presented in (Ridaura,

2014, Martins, 2015). Figure 8, 9, 10 and 11 show acceptable agreement between the simulation

results with the test data for the fan, LPC and HPC operating lines for steady-state test at Sea-Level-

Static (SIS) condition from idle to full throttle setting. However, part load SFC (shown in Figure 11)

depends primarily on compressors’ efficiency for which, the operating points from the CFM56-3

correlation test report do not match with the model due to the use of generic component maps. In

order to make the model match with test results slight modifications have been done on HPC’s

efficiency map. Results for SFC are illustrated in Figure 11. The SFC starts decreases from 100 kN

to 75kN due to the increase in HPC efficiency. While between 75kN to 65kN of thrust, the HPC

operation point hits its maximum efficiency which is 87 % on the HPC map. Below 65kN of thrust,

the VBVs and VSVs start to act which affect the SFC severely, decreasing the SFC for 65kN to 55kN

of thrust and increasing it below 55kN of thrust.

Figure 8 FAN operating line (idle to full

throttle at SIS condition)

Figure 9 LPC operating line (idle to full

throttle at SIS condition)

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Figure 10 HPC operating line (idle to full

throttle at SIS condition)

Figure 11 Specific Fuel consumption (idle to

full throttle at SIS condition)

POWER TRANSFER RESULTS

In this section, performance improvements due to power transfer between shafts for a turbofan at

taxiing and flight idle maneuver are presented. In order to investigate the idea of power exchange

between shafts, considering losses in gearboxes, electrical machines and converters, overall

efficiency of the Electric Power Transfer (EPT) is assumed to be 90%.

Results for ground and flight idle scenarios are presented for a CFM56-3 engine with 100 kN of

thrust at SLS condition. The derived values of 𝑃𝐹𝐴𝑁, 𝑃𝐿𝑃𝐶 , 𝑃𝐻𝑃𝐶 , 𝑃𝐻𝑃𝑇 and 𝑃𝐿𝑃𝑇 from the engine

model are fed to the LP and HP shaft modules together with the amount of electric power transfer

between them in accordance with the dynamics given in equations 1 to 2.

Taxiing

For taxi maneuver, a 10% of engine thrust is required based on ICAO values in (Chati and

Balakrishnan, 2013). For the purpose of comparison a controller keeps the thrust at its minimum value

as it is shown in Figure 12. Since the fan provides most of the thrust, by electric power transfer, LP

speed remains unchanged, while HP speed increases. Plotted in Figure 13, the compressors’ SM also

increases as expected by increase in HPC speed by 5% and 2.5% for the LPC and HPC respectively.

Moreover, the fuel rate at taxi reduces by approximately 3%.

Figure 12. Thrust and shaft speeds (taxi)

Figure 13. Available SM and Fuel rate (taxi)

Flight Idle

During descent, flight idle mode is selected where the engine needs to operate at a high enough

power to keep the bleeding pressure higher than a minimum for its application in the ECS. In order

to keep the engine operation within its limits while allowing immediate recovery of thrust during

power transfer, the controller must maintain HP speed at its default minimum idle setting which has

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also been assumed to be 70% of its nominal value. 𝑝30 limit controller protects against the supply

pressure for ECS (Linke-Diesinger, 2010) which has been assumed to be 210kPa for this engine.

Results for flight idle at 10kft and 30kft are presented in Figure 14 to 17.

At 10kft, up to 190kW power transfer, by keeping the HP speed at its minimum permissible value

the fuel rate decreases by 26%, (shown in Figure 14 and 15). Since the power is extracted from the

LP shaft its speed decreases by 25% which reduces the thrust more or less by the same value. The 𝑝30

level is more than its minimum value at 10kft and the minimum controller switches from the HP shaft

speed to the LP one at 190 kW. Indicated in Figure 15, SM for the LPC and HPC increase by 13%

and 10% respectively.

At 30kft, 𝑝30 limit controller is acting within 200kW of power transfer. Results over 200kW of

electric power transfer is not valid since LPT cannot provide enough power for electric power transfer.

The small deviation from 210kPa shown in Figure 16 is due to the transient of the p30 controller

during increasing the amount of electric power transfer. By 200kW electric power transfer, LP speed

drops by 30% while HP speed remains at its existing level to ensure the pressure of 210kPa at HPC

exit. Shown in Figure 17, available SM for the LPC and HPC increases by 13% and 14% respectively,

and the fuel rate decreases by almost 31%.

Figure 14. 𝒑𝟑𝟎 and shaft speeds (flight

idle-10kft)

Figure 15. Available SM and fuel rate (flight

idle-10kft)

Figure 16. 𝒑𝟑𝟎 and shaft speeds (flight

idle-30kft)

Figure 17. Available SM and fuel rate (flight

idle-30kft)

Cruise condition

In order to have a proper assessment of electric power transfer for cruise condition, a controller

has been implemented to keep the thrust at its existing value without the electric power transfer

system. Shown in Figure 18, the assumed value of thrust is 26kN for CFM56-3 (Chati and

Balakrishnan, 2013) which is aligned with propulsive efficiency for the aircraft at cruise condition.

The engine efficiency is at its maximum level with the existing configuration for cruise condition

with 26kN of thrust. Therefore, electric power transfer is not expected to reduce the fuel rate, however,

it can increase SM of compressors. Figure 19 highlights the fact that any power transfer increases the

9

fuel rate around the existing design point. For ±300kW electric power transfer, the amount of increase

in fuel rate is around 4.5% and the amount of increase in SM for the LPC and HPC are approximately

5% and 8% respectively. Therefore, electric power transfer at cruise condition is not beneficial unless

for increasing SM.

Figure 18. Thrust and shaft speeds

(cruise)

Figure 19. Available SM and fuel rate (cruise)

CONCLUSIONS

This paper presented a model for the CFM56-3 engine with VBVs and VSVs involved. Validation

with test results indicated acceptable level of accuracy for the engine model although with slightly

modified generic engine maps from GASTURB. Furthermore, the paper provided preliminary results

for the impact of electrical power transfer for the CFM56-3 engine. Power was proposed to be

circulated between the electrical machines, connected to the engine shafts, through the single DC bus

architecture with the assumed efficiency of 90%. Results showed that, power transfer from the LP to

the HP shaft at low compressor speeds is highly desirable both in terms of available SM and

efficiency. These are the outcomes of increased HPC speed operating at higher efficiency and away

from surge line. For Taxi 3% and for flight idle more than 25% of fuel rate reductions are achieved

by the electric power transfer system. However, electric power transfer at cruise condition increases

the fuel rate which makes this idea futile for cruise unless for increasing SM by positive electric

power transfer. The results in this paper motivate further analysis into the idea of electric power

transfer for turbofans in future.

ACKNOWLEDGEMENTS

This project has received funding from the Clean Sky 2 Joint Undertaking under the European

Union’s Horizon 2020 research and innovation programme under grant agreement No 807081.

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