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ISABE-2017-22643 (RAeS Review, Subject Area: Engineering Sciences), Sept. 3-8, 2017, Manchester, UK

(DRAFT)

STUDY OF A CONCEPTUAL DESIGN FOR COOLED COOLING AIR IN A PRESWIRL CAVITY

Zixiang Sun and John W. Chew [email protected]

Thermo-Fluid Systems UTCDept. of Mech. Eng. Sciences

University of SurreyGuildford, Surrey, GU2 7XH

UK

ABSTRACTTo achieve enhanced cooling of hot components in the high pressure (HP) section of an aeroengine,

application of cooled cooling air (CCA) has been proposed. Here a “two row preswirl feed” arrangement is considered to accommodate the CCA, together with the uncooled cooling air (UCA) in high pressure turbine (HPT) preswirl cavity. The CCA and UCA inflows are introduced into the preswirl cavity at two different radii. Most of the cooling air leaves the preswirl cavity from the receiver holes. To assess the CCA behavior in the preswirl cavity, a definition of feeding effectiveness is introduced based on the relative total temperature at the exit of the receiver hole. The CFD investigation for the preswirl cavity was conducted in a systematic way by altering both the radial position of the receiver hole and inflows of the CCA and UCA, while keeping other conditions unchanged. It was found that the feeding effectiveness increases as the radial position of the receiver hole decreases. An optimal feeding effectiveness close to a minimum mixing condition was achieved by adjusting the CCA and UCA inflows. Unsteady CFD investigations gave a similar prediction for the overall performance of the CCA in the preswirl cavity, but with a lower feeding effectiveness compared with its steady CFD counterpart. The reduction in the feeding effectiveness was attributed to an enhanced mixing from the discrete CCA and UCA inflows and associated unsteady effects.

Key Word: Cooled cooling air (CCA); Enhanced cooling; Conceptual Design.

Z. Sun and J. W. Chew 1 Copyright © 2017 Rolls-Royce plc


NOMENCLATURECp specific molar heat capacity at constant pressuredh hydraulic diameter at inlet = 4*Area/Wetted-perimeterI Rothalpy

mass flow ratePCR pitch circle radius pout outlet pressure at receiver holeRebulk inflow Reynolds number = Ubulk dh/Re rotational Reynolds number = ro

2/ r radiusr1,2 entrance and exit of receiver hole, respectivelyro outer radius of the cavityT temperatureTref reference temperatureT*

rot rotary stagnation temperatureTt total temperatureTtrel relative total temperatureUbulk bulk velocity at inletV swirl velocityy+ dimensionless wall distance = (w/)0.5yp/yp wall distance

Greekrel feeding effectiveness, Equation 3 dynamic viscosity angular velocity of rotor or disc densityw wall shear stress

Subscriptsbaseline baseline test caseCCA cooled cooling air inflowFull full mixingUCA uncooled cooling air inflowi inner radiusin_seal inner labyrinth sealm middle radiusmin minimum or minimum mixingRecHol receiver holeSum total mass flow rate through the cavityo outer radiusw wall

1. INTRODUCTIONUse of cooled cooling air to achieve enhanced cooling of hot components in aeroengine applications has

been considered in recent years. A CCA concept was investigated as part of the “active core” research under the European Framework 6 collaborative project NEW Aero engine Core concepts (NEWAC, 2006~2011). An early publication about the CCA option in the “active core” task under the NEWAC consortium was given by Wilfert et al. in 2007 [1]. Relevant progress on the CCA work was reported by Sturm [2] and Ebert et al. [3] at the European Workshop on New Aero Engine Concepts held in Munich, Germany, 2010. A further description of the CCA study in the “active core” research was published by Bock et al. in 2008 [4]. Other ideas for aeroengine applications of CCA were claimed in patents, such as European patent EP2275656A2 by Chir and Edwards in 2011 [5] and US patent 6612114 by Hermann in 2003 [6] amongst others.

Most of the previous CCA researches have been limited to conceptual studies and assessments of the CCA technology. In the present study, a more detailed aspect of CCA application is considered. In particular, a “two row preswirl feed” arrangement to accommodate the CCA in a HPT preswirl cavity is investigated. The temperature of the air delivered to the blade cooling feed is represented through a feeding effectiveness, and it is shown that this depends on the preswirl arrangement.



Fig. 1 Illustration of a Cooled Cooling Air Application

The proposed cooled cooling air application is illustrated in Figure 1. A portion of hot air is extracted from the main gas path of the high pressure compressor (HPC), and cooled by the bypass air through a heat exchanger. The cooled air is then fed into the secondary air system through the HPC drive cone cavity and HPT preswirl cavity to provide an enhanced cooling for hot components of the engine, such as the HPC rear cone, HPT blade and disc rim, etc.

A possible two row preswirl feed arrangement of the CCA and UCA streams is described in the next section. The CFD model used, the scope of the investigation, definition of the feeding effectiveness, results and discussion are then presented in the following sections.

2. TWO ROW PRESWIRL FEED ARRANGEMENTA sectional view of a “two row preswirl feed” to the HPT preswirl cavity for the CCA application is shown

in Figure 2. The geometry is based on an industrial configuration, and was modified to accommodate the CCA. It can be seen that the preswirl cavity is bounded by stationary structures on the left and the rotating disc on the right. The receiver holes on the right also rotate. The boundaries separating the stationary and rotating walls are the inner and outer labyrinth seals, located at the inner and outer radii of the preswirl cavity, respectively. In addition to the seal inflow and outflow, there are two more inflows and one further outflow in the cavity. The CCA and UCA inflows were arranged at two different radii in the stationary structures on the left, with the CCA being placed at a slightly lower pitch circle radius (PCR). Such an arrangement for the CCA is thus called “two row preswirl feed”. Most of the cooling air outflow leaves the cavity through the receiver holes on the right.

Fig. 2 Two row preswirl feed

It may be noted that the seal flows are significant and have an important influence on the temperature of the cooling air delivered to the receiver holes. Mixing of the three inlet streams will also affect the cooling air delivery temperature.



3. CFD MODEL AND NUMERICAL ISSUESThe CFD model for the engine representative preswirl cavity employed in the present study is shown in

Figure 3. It is a 2.73° sector assuming circumferential periodicity, with one discrete receiver hole. A contraction was added to the exit of the receiver hole to eliminate reverse flow there. The inner and outer seals plus the CCA and UCA inflows were approximated with annular slots. All the walls are axi-symmetric except the receiver hole. The mesh was generated using ICEM CFD software. All the mesh cells are of hexahedral type, with 634,020 cells. The mesh resolution follows previous mesh-dependency investigations and is similar to the authors’ two previous publications [7, 8] regarding preswirl cavities. The mesh was thus considered adequate. The dimensionless wall distance y+ obtained was generally kept between 50 and 100.

Fig. 3 The Preswirl Cavity CFD Model (2.73° Sector)

The boundary condition types of the CFD model are also shown in Figure 3. All the three inlets are specified as subsonic inlet with assigned mass flow, total temperatures and swirl velocities. In the figure, , Tt

and V denote mass flow rate, total temperature and swirl velocity, respectively. The outer labyrinth seal and receiver hole are treated as a subsonic outlet with specified flow ( ) and an outlet with assigned pressure (

), respectively. The mass flow rate ( ) at the receiver hole is determined by the CFD solution, and may slightly deviate from a mass balance analysis. The right rotor walls are under rotation with an angular speed of . All the walls were assumed no-slip and adiabatic in the CFD simulations.

(4-a) at PCRo (4-b) at PCRm (4-c) at PCRi Fig. 4 The Preswirl Box CFD Models (2.73° Sector)

Three simplified models were further constructed for a comparative study using the 2.73° sector with a discrete receiver hole. These are shown in Figure 4. It can be seen that most geometrical features of the preswirl cavity, such as inlets and outlets were kept in these simplified “box” models. In addition, all the three box models are identical apart from the radial position of the receiver hole. The pitch circle radii (PCR) of the receiver hole for the three box models are at the outer, middle and inner radii PCR o, PCRm and PCRi, respectively. Contractions were again added to the exit of the receiver hole to eliminate reverse flow there. A box model with a discrete preswirl nozzle rather than a slot was also considered and is described later.



The CFD simulations were conducted using the standard k- turbulence model and wall functions. Turbulence modeling is fully justified as the Reynolds number of the flow is high. Taking the baseline test case for example, its rotational Reynolds number is equal to Re=ro

2/=3.13x107, where and stand for density and viscosity, respectively, and ro denotes the outer radius of the preswirl cavity. The through-flow Reynolds numbers are equal to Rebulk=Vbulkdh/ =1.10x107, 7.28x104 and 4.96x104 for the UCA, CCA and the inner labyrinth seal inflows, where Vbulk and dh denote the bulk velocity of the inflows and the hydraulic diameter of the inlets, respectively.

The CFD solver employed in the investigation was the Rolls-Royce Hydra code. The numerical scheme used is of 2nd order accuracy. Both steady and unsteady CFD simulations were performed.

Convergence of all CFD simulations was monitored according to the standard practice. For example, in the preswirl cavity model under the baseline condition, the overall balance of mass, angular momentum and total enthalpy are 0.1%, 2% and 3% in terms of total mass inflow rate, rotor shaft torque and rotor windage, respectively.

4. TEST CASE MATRIXSeven test conditions including the baseline condition, were employed in the present parametric study. The

baseline test case represents a maximum take-off (MTO) condition. The remaining six test conditions were obtained from the baseline by altering the CCA and UCA inflows systematically, while keeping all other boundary conditions unchanged. In fact, only three variables of the boundary conditions for the CCA and UCA inflows were changed. They are the mass flow rates for both the CCA and UCA, plus the total temperature of the CCA. The change of the above three items was further constrained by the total mass and inflow energy requirements. For all the seven test cases, the total mass inflows and total energy (enthalpy) influx were kept unchanged.

(1)

(2)

where Cp denotes specific heat at constant pressure, which was also assumed constant in the present investigation.

Fig. 5 Changes of Boundary Conditions

A summary of alterations in boundary conditions for CCA and UCA is given in Table 1. The corresponding graphic representation of the boundary condition changes is shown in Figure 5. It can be seen that as the mass flow from CCA decreases, the mass flow from UCA increases accordingly to keep the total mass inflow unchanged. The total temperature of the UCA was kept constant for all the test cases. As a result, the total temperature of the CCA has to be reduced to keep the total enthalpy influx from inlets constant when the mass flow from CCA decreases. In addition, the swirl velocities for all the inflows were kept unchanged.



Table 1 Test Case Matrix

Test CaseCCA (V=Const.) UCA (Tt, V=Const.)

Tt/Tref

Baseline 1.0 0.280 2.462 0.521No. 1 1.1 0.308 2.539 0.493No. 2 0.9 0.252 2.367 0.549No. 3 0.8 0.224 2.248 0.577No. 4 0.75 0.210 2.177 0.591No. 5 0.7 0.196 2.096 0.605No. 6 0.5 0.140 1.608 0.661

Where mSum (=mCCA+mUCA+mIn-Seal) and Tref denote the total mass flow through the cavity and a reference temperature of the engine.

5. DEFINITION OF FEEDING EFFECTIVENESSTo facilitate the assessment of the CCA behavior in the preswirl cavity, a definition of feeding effectiveness

is given as follows:

(3)

where rel stands for the feeding effectiveness based on relative total temperature. T trel_RecHole denotes the relative total temperature at the receiver hole exit obtained in the CFD modelling. T trel_RecHole@Min_mixing and Ttrel_RecHole@Full_mixing represent the relative total temperatures at receiver hole obtained by an analytical estimation at the idealized minimum and full mixing situations, respectively.

An illustration of the idealized minimum and maximum mixing assumptions is given in Figure 6. In the idealized full mixing situation, all the inflows are assumed to be fully mixed in the preswirl cavity before entering the receiver hole and the outer labyrinth seal. In the minimum mixing condition, it is assumed that the inflows from CCA and the inner labyrinth seal enter the receiver hole without mixing with the UCA. Part of the UCA joins the CCA and the inflow from the inner labyrinth seal to go through the receiver hole, while rest of the UCA is isolated from other inflows and leaves the cavity from the outer labyrinth seal. It may be noted that fully mixed preswirl chamber models have previously been used successfully in convectional systems [9].

(6-a) Full Mixing (6-b) Minimum Mixing

Fig. 6 Illustration of Full and Minimum Mixings

Estimation of the relative total temperature at the receiver hole for the idealised full and minimum mixings is executed with the following two steps. Considering the full mixing case, 1) at position r 1, the entrance of the



receiver hole (Figure 6-a), the total temperature is obtained from a mass weighted average of total enthalpy influx from all the three inflows.

(4)

Similarly, the absolute swirl velocity at r1 is obtained from a mass weighted average of total angular momentum influx from all the three inflows:

(5)

2) At position r2, the exit of the receiver hole, the rothalpy there is assumed to be conserved through the receiver hole from its entrance at position r1.

(6)

where rothalpy I is defined as follows

(7)

The relative total temperature at position r2 can then be obtained from the relationship between temperature and rothalpy.

(8)

When needed, the corresponding absolute total temperature can be obtained by further assuming the absolute swirl velocity there is at solid body rotation.

(9)

Estimation of the relative total temperature at the receiver hole for the idealised minimum mixing case follows similar procedures. Firstly, the amount of mass flow rate contributing to the receiver hole flow from the UCA can be determined using a mass balance analysis (Figure 6-b).

(10)

where stands for the minimum mass flow contribution from UCA to the receiver hole, and the subscripts RecHole, in_seal and CCA denote the flows from the receiver hole, the inner labyrinth seal and CCA, respectively.

Then, the remaining procedures for the full mixing case are applied. 1) At first at position r 1, the entrance of the receiver hole, the total temperature is estimated from a mass weighted average of total enthalpy influx from the contributing inflows.

(11)

Similarly, the absolute swirl velocity is estimated from a mass weighted average of total angular momentum influx from the inflows.



(12)

2) At position r2, the exit of the receiver hole, the relative total temperature can be obtained from the assumption of rothalpy conservation through the receiver hole using Equations 6 to 9.

Note that the windage contribution to the relative total temperature was not taken into account in the present estimation, as the windage and resultant temperature rise in the cavity were expected to be small due to high swirl velocity from the inflows. The CFD predictions did include the windage contribution and confirmed that windage in the cavity was small compared with the difference in enthalpy influx of the two coolant streams.

2.70

2.72

2.74

2.76

2.78

2.80

2.82

2.84

70 75 80 85 90 95 100 105 110Ratio of CCA Mass Flow mCCA / mCCA_Baseline %

Norm

alis

ed R

elat

ive

Tota

l Tem

pera

ture

Ttre

l/Tre

f Ttrel_RecHole@Full-MixingTtrel_RecHole@Min-Mixing

Fig. 7 Relative Total Temperatures at the Exit of the Receiver Hole for the Preswirl Cavity with Idealised Full and Minimum Mixing (Analytical Estimations)

The relative total temperature at the receiver hole obtained for the preswirl cavity at the idealised full and minimum mixing situations is shown in Figure 7. It can be seen that the relative total temperature decreases as the normalized mass flow rate from CCA is reduced. The difference of the relative total temperature between the full and minimum mixings is almost unchanged.

6. STEADY CFD RESULTSTypical flow patterns obtained for the engine representative geometry are shown in Figure 8. The CFD

model used is the 2.73° sector of preswirl cavity with discrete receiver hole (Figure 3). The results are for the baseline condition. The contour plots show the paths of the CCA and UCA jet inflows. The higher swirl of the inflows reduces as the diffusion and mixing processes take effect inside the cavity. Similar phenomena can also be observed from the relative total temperature contours. The cold CCA and hotter UCA jets gradually diminish in the cavity. In addition, stratifications of the flow and temperature fields are also visible. The swirl ratio is apparently higher in the receiver hole and the outer part of the cavity. Higher relative total temperature is noticeable in the outer part of the cavity. 3D features of the flow are obvious. Non-uniform distributions of the flow in the circumferential direction can be clearly identified in the region of the receiver hole.

A comparison of swirl ratio and relative total temperature for the box models with the receiver hole at different radial positions is shown in Figure 9. These results are again for the baseline condition. The contours were plotted using the same colour scale as in Figure 8. It can be seen that the flow and thermal mixing in the cavity are considerably different for the three box models. When the receiver hole moves radially inwards, the penetration depths of the CCA and UCA jets become longer and the stratification of the flow and relative total temperature in the cavity is strengthened. All these indicate that the mixing effect becomes weaker and



consequently the relative total temperature obtained at the exit of receiver hole is reduced when the pitch circle radius (PCR) of the receiver hole decreases.

Vr Ttrel/Tref

0

1.2

2.3

3.3

(8-a) Swirl Ratio (8-b) Relative Total Temperature

Fig. 8 Flow Features of Two row preswirl feed, 2.73° Sector Model for Preswirl Cavity with Discrete Receiver Hole, Baseline Test Case

A graphical comparison of the feeding effectiveness for the three box models under the baseline condition is given in Figure 10. The corresponding result from the 2.73° sector of the engine representative model is also inserted. The horizontal axis is the pitch circle radius (PCR) of the receiver hole normalized by the cavity outer radius ro. The two vertical coordinates are the relative total temperature (denoted by red diamonds) and feeding effectiveness (represented by solid blue cycles). The figure clearly shows that the relative total temperature obtained at the exit of receiver hole decreases when the pitch circle radius PCR of the receiver hole is reduced. When transformed into feeding effectiveness, it can be seen that the feeding effectiveness increases as the PCR decreases. The negative feeding effectiveness at the low-right corner in the figure obtained for the engine representative cavity model is due to the windage effect being taken into account in the CFD simulations, while in the analytical estimations for the full and minimum mixing cases that term was omitted for simplification. The CFD simulations indicate that the higher the pitch circle radius of the discrete receiver hole, the greater the rotor walls’ windage for the HPT preswirl cavity, although the windage itself is small compared with changes in total enthalpy. The windage heating from the rotating walls inside the cavity was estimated to increase the overall cavity temperature by 0.1~0.6%Tref at the present MTO condition, dependent on the radial position of the discrete receiver hole.

Vr

Ttrel/Tref

PCRi

PCRm

PCRo

0

1.2

2.3

3.3

Fig. 9 Effect of the Radial Position of the Receiver Hole on Two row preswirl feed, 2.73° Sector Models, Baseline Test Case

The effect of inflow conditions on feeding effectiveness was also investigated by systematically altering the CCA and UCA inflows, as given in Table 1. A typical result of the parametric study is shown in Figure 11. The CFD model in question is the box model with the receiver hole at the inner radius PCR i (Figure 4-c). The



corresponding relative total temperatures estimated at the exit of the receiver hole for the full and minimum mixing cases are also given in the figure. It can be seen that the relative total temperature at the exit of receiver hole obtained by the sector model exhibited a U-shape in the present investigation. It approached closely to its minimum mixing counterpart around some 80% to 90% of the baseline CCA mass inflow. At that point, the feeding effectiveness reached its maximum, as shown in the figure.

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0.88 0.9 0.92 0.94 0.96 0.98 1Normalised Pitch Circle Radius of Receiving Hole PCR/ro

Nor

mal

ised

Rel

ativ

e To

tal T

emp.

Ttre

l/Tre

f

-40

-20

0

20

40

60

80

100

Feed

ing

Effe

ctiv

enes

s

rel (

%)

Trel@RecHole, CFDFeeding Effectiveness h_rel

Fig. 10 Comparison of CFD Results for the Two row preswirl feed, 2.73° Sector Models, Baseline Test Case

Fig. 11 Effect of Inflow Conditions on Two row preswirl feed, 2.73° Sector Model with Discrete Receiver Hole at the Inner Radius PCRi

Further analysis was conducted to understand the feeding effectiveness drop at low CCA flows. It was found that the penetration length of CCA jet decreases when the CCA mass inflow is reduced. In addition, the trajectory of the CCA jet is deflected radially outwards away from its axial direction. It was thus understood that the drop in effectiveness is related to the behavior of the CCA jet. The influence of UCA jet may also play a part as its strength is enhanced when the CCA mass inflow decreases.

The two row preswirl feed was also investigated using a slot approximation for all the inlets and outlets. In other words, the discrete receiver hole was also modeled by an annular slot in the further numerical investigations. The feeding effectiveness thus obtained was greater than its counterpart with the discrete receiver hole. Hence, caution is needed in use of a slot approximation for receiver holes, and there may be some sensitivity to the number and size of the receiver holes.



7. UNSTEADY CFD RESULTSAn unsteady CFD investigation was conducted in order to gauge the significance of the effects of the flow

unsteadiness and discrete inflows. The geometry of the box-1 model with the discrete receiver hole at the inner pitch circle radius PCRi was chosen with appropriate modifications. The resultant CFD model is shown in Figure 12. It can be seen that the model consists of two domains, separated by a sliding plane between the discrete CCA inlet and the discrete receiver hole. The outer part contains the discrete CCA and UCA inflows in addition to the outer labyrinth seal. The inner part keeps the discrete receiver hole and the inner labyrinth seal. The outer domain was set to be in the absolute reference frame, and the inner domain was kept in the relative rotating reference frame. Communication of information between the two domains is realized through the sliding plane.

Fig. 12 The Preswirl Box CFD Model for Unsteady CFD Simulation Using Sliding Plane (2.73° Sector Model)

0

10

20

30

40

50

60

70

80

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70 75 80 85 90 95 100 105 110Ratio of CCA Mass Flow mCCA / mCCA_Baseline (%)

Feed

ing

Effe

ctiv

enes

s tre

l (%

)

Feeding Effectiveness hrel % (Box-1@PCRi, Steady CFD) hrel % (Box-4@PCRi, Unsteady CFD)

PCRi

Fig. 13 Comparison between Steady and Unsteady CFD Simulations for the Two row preswirl feed, 2.73° Sector Box Model with Discrete Receiver Hole at the Inner Radius PCR i

The unsteady CFD investigation employed an implicit time-stepping algorithm, and was conducted under the similar specifications of inflow and outflow boundary conditions to the steady solutions. The time step was set to be equal to 1/7920 of the disc rotating frequency, which is equivalent to one circumferential mesh spacing per time step. Turbulence was again simulated using the k- turbulence model together with wall functions. Convergence to a periodic solution was achieved after a few rotor revolutions. The unsteady CFD solutions were then time-averaged over a further one revolution. The time-averaged unsteady CFD results were compared with their steady CFD counterparts accordingly. It was found that the relative total temperatures at the exit of



receiver hole obtained for the unsteady CFD simulations were higher than their steady CFD counterparts. When transformed into the feeding effectiveness, it was revealed that the feeding effectiveness obtained from the unsteady CFD simulations is lower than its steady CFD counterpart. This is shown in Figure 13. This effect is attributed to enhanced mixing from the discrete inflows and associated unsteady effects.

8. CONCLUSIONSDesign of a preswirled delivering system incorporating cooled cooling air CCA has been considered. A

“two row preswirl feed” arrangement was proposed. The preswirl chamber is supplied with separate streams of CCA, UCA and sealing air, and the degree of mixing of these streams affects the system performance. A definition of feeding effectiveness was given. The steady CFD shows that the feeding effectiveness increases as the pitch circle radius PCR of receiver hole decreases. An optimal feeding effectiveness close to the minimum mixing limit was further achieved by adjusting CCA and UCA flows. Compared with the steady CFD results, similar trends were obtained from the unsteady CFD simulations but with a lower feeding effectiveness prediction. This was attributed to induced mixing from the discrete inflows and associated unsteady effect.

ACKNOWLEDGMENTSFunding from the Department of Trade and Industry (DTI) and Rolls-Royce plc is gratefully acknowledged.

The project was coordinated by Ivan Popovic and John Irving from Rolls-Royce plc. Sincere thanks also go to Guy D. Snowsill from Rolls-Royce plc for his comments and discussions.

REFERENCES[1] Wilfert, G., Sieber, J., Rolt, A., Baker, N., Touyeras A., and Colantuoni, S., 2007, “New Environmental

Friendly Aero Engine Core Concepts”, Proc. of 18th International Symposium on Airbreathing Engines (ISABE), Beijing, China, September 2007, Paper No. ISABE-2007-1120.

[2] Sturm, W., 2010, ”Active Core (an Overview)“, European Workshop on New Aero Engine Concepts, Munich, Germany, 30 June – 1 July 2010, http://www.newac.eu/88.0.html.

[3] Ebert, E., Klingels, H. and Storm, P.,2010, ”Concept Study on an Advanced Cooling Air Cooling System“, European Workshop on New Aero Engine Concepts, Munich, Germany, 30 June – 1 July 2010, http://www.newac.eu/88.0.html.

[4] Bock, S., Horn, W., and Sieber , J., 2008, “Active Core” – a Key Technology For More Environmentally Friendly Aero Engines Being Investigated under the Newac Program”, Proc. 26th International Congress of the Aeronautical Sciences , 14-19 September 2008, Anchorage, Alaska

[5] Chir, A. P. and Edwards, H. L., 2011, “System for cooling-air in a gas turbine engine”, European patent EP 2 275 656 A2, 2011.

[6] Hermann, K. 2003 “Cooling air system for gas turbine”, United States patent 6612114, 2003.

[7] Sun, Z., Chew, J.W., Hills, N.J., Lewis, L. and Mabilat, C., 2012, “Coupled Aerothermomechanical Simulation for a Turbine Disk Through a Full Transient Cycle”, Trans. ASME Journal of Turbomachinery, v 134, n 1, p 011014 (11 pp.), Jan. 2012, see also Proc. ASME Turbo Expo 2010, June 14-18, 2010, Glasgow, UK, Paper no. GT2010-22673, pp. 1-9.

[8] Sun, Z., Chew, J.W., Hills, N.J., Barnes, C. and Valencia, A., 2012, “3D Coupled Fluid-Solid Thermal Simulation of a Turbine Disc Through a Transient Cycle”, Proc. ASME Turbo Expo 2012, June 11-15, 2012, Copenhagen, Denmark, Paper No. GT2012-68430.

[9] Chew, J.W., Hills, N.J., Khalatov, S., Scanlon, T. and Turner, A.B., 2003, “Measurement and Analysis of Flow in a Pre-swirled Cooling Air Delivery System”, Proc. ASME Turbo Expo 2003, June 16-19, 2003, Atlanta, Georgia, USA, Paper No. GT2003-38084.


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