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Research Papers

The Effect of External Casing Impingement Cooling Manifold Standoff Distance on Casing Contraction for Thermal Control of Blade Tip Clearance

[+] Author and Article Information
Myeonggeun Choi

Department of Engineering Science,
University of Oxford,
Oxford OX1 3PJ, UK
e-mail: myeonggeun.choi@eng.ox.ac.uk

David R. H. Gillespie

Department of Engineering Science,
University of Oxford,
Oxford OX1 3PJ, UK
e-mail: david.gillespie@eng.ox.ac.uk

Leo V. Lewis

Rolls-Royce plc,
P.O. Box 31,
Derby DE24 8BJ, UK
e-mail: leo.lewis@rolls-royce.com

Contributed by the International Gas Turbine Institute (IGTI) of ASME for publication in the JOURNAL OF TURBOMACHINERY. Manuscript received August 16, 2017; final manuscript received October 6, 2017; published online November 29, 2017. Editor: Kenneth Hall.

J. Turbomach 140(2), 021005 (Nov 29, 2017) (13 pages) Paper No: TURBO-17-1117; doi: 10.1115/1.4038280 History: Received August 16, 2017; Revised October 06, 2017

Thermal closure of the engine casing is widely used to minimize undesirable blade tip leakage flows thus improving jet engine performance. This may be achieved using an impingement cooling scheme on the external casing wall, provided by manifolds attached to the outside of the engine. The assembly tolerance of these components leads to variation in the standoff distance between the manifold and the casing, and its effects on casing contraction must be understood to allow build tolerance to be specified. For cooling arrangements with promising performance, the variation in closure with standoff distance of z/d = 1–6 were investigated through a mixture of extensive numerical modeling and experimental validation. A cooling manifold, typical of that adopted by several engine companies, incorporating three different arrays of short cooling holes (chosen from previous study by Choi et al. (2016, “The Relative Performance of External Casing Impingement Cooling Arrangements for Thermal Control of Blade Tip Clearance,” ASME J. Turbomach., 138(3), p. 031005.)) and thermal control dummy flanges were considered. Typical contractions of 0.5–2.2 mm are achieved from the 0.02–0.35 kg/s of the current casing cooling flows. The variation in heat transfer coefficient observed with standoff distance is much lower for the sparse array investigated compared to previous designs employing arrays typical of blade cooling configurations. The reason for this is explained through interrogation of the local flow field and resultant heat transfer coefficient. This implies that acceptable control of the circumferential uniformity of case cooling can be achieved with relatively large assembly tolerance of the manifold relative to the casing.

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Figures

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Fig. 5

Computational domain for standalone external casing heat transfer prediction

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Fig. 4

Cooling arrangements selected for manifold offset parametric study: baseline H2, fillet cooling H2C, and combination cooling H5CI

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Fig. 8

(a) Schematic diagrams of the main components of the test rig and (b) a photograph of the test section and typical liquid crystal activation under H2 cooling array

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Fig. 3

A typical turbine casing cooling configuration with impingement through a manifold onto the casing wall

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Fig. 2

A typical radial cross section of turbine casing cooling arrangement, adopted from Lewis and Bacic [4]

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Fig. 6

Comparison of H2 solutions showing near grid independent solutions achieved for y+ < 1.5

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Fig. 7

Idealized representation of the casing FE model for thermomechanical analysis

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Fig. 12

Local htc/htcref (=H2 z/d=4.peak.realizablek–ε.OP1) distributions for H2 z/dbase = 1–6 scheme at OP1 (CFD)

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Fig. 13

Spanwise-averaged htc/htcref (=H2 z/d=4.peak.span.realizablekε.OP1) distributions for H2 z/dbase = 1–6 scheme (CFD)

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Fig. 1

Potential tip clearance improvement induced by turbine casing thermal control system

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Fig. 16

Spanwise-averaged htc/htcref (=H2 z/d=4.peak.span.realizablekε.OP1) distributions for H2C z/dbase = 1–6 scheme (CFD)

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Fig. 9

Selected streamlines colored by vel/velmax at OP1

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Fig. 10

Local htc/htcref (=H2 z/d=4.peak.span.realizablek–ε.OP1) distributions for H2 z/dbase = 1–6, experimental case Re = 6000 (EXP), equivalent to OP1

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Fig. 11

Spanwise averaged htc/htcref (=H2 z/d=4.peak.span.realizablek–ε.OP1) distributions for H2 z/dbase = 1–6, experimental case Re = 6000 (EXP), equivalent to OP1

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Fig. 14

(a) Patch area-averaged htc/htcref (=H2 z/d=4.patch.realizablekε.OP1) at OP1 and (b) Relative performance of overall area-averaged htc/htcref (=H2 z/d=4.avg.realizablek–ε.OP1) at OP1 and OP2 (CFD)—including experimental results

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Fig. 15

Local htc/htcref (=H2 z/d=4.peak.realizablekε.OP1) distributions for H2C z/dbase = 1–6 scheme at OP1 (CFD)

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Fig. 19

Predicted casing temperature (K) and thermal radial displacement (r, mm) for (a) H2 z/dbase = 4, (b) H2C z/dbase = 4, (c) H5CI z/dbase = 4 at OP1, and (d) typical distributions along the casing length for H2 case

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Fig. 20

Predicted casing temperature (K) and closure (mm) versus coolant mass flow rate, at liner attachment point, for (a) H2 z/dbase = 1–6, (b) H2C z/dbase = 1–6, and (c) H5CI z/dbase = 1–6 scheme

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Fig. 21

Relative performance of casing closure at the liner attachment point by internal hooks, compared to baseline H2 z/d = 4 scheme at OP1

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Fig. 17

Local htc/htcref (=  H2 z/d=4.peak.realizablek–ε.OP1) distributions for H5CI z/dbase = 1–6 scheme at OP1 (CFD)

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Fig. 18

Spanwise-averaged htc/htcref (=H2 z/d=4.peak.span.realizablek–ε.OP1) distributions for H5CI z/dbase = 1–6 scheme (CFD)

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