esdu inlets characteristics

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ESDU 86002

Endorsed byThe Royal Aeronautical Society

Drag and pressure recoverycharacteristics of auxiliary air inlets at

subsonic speeds

Issued April 1986With Amendments A to D

June 2004Supersedes ESDU 66029

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ESDU 86002ESDU DATA ITEMS

Data Items provide validated information in engineering design and analysis for use by, or under the supervisionof, professionally qualified engineers. The data are founded on an evaluation of all the relevant information, bothpublished and unpublished, and are invariably supported by original work of ESDU staff engineers or consultants.The whole process is subject to independent review for which crucial support is provided by industrial companies,government research laboratories, universities and others from around the world through the participation of someof their leading experts on ESDU Technical Committees. This process ensures that the results of much valuablework (theoretical, experimental and operational), which may not be widely available or in a readily usable form, canbe communicated concisely and accurately to the engineering community.

We are constantly striving to develop new work and review data already issued. Any comments arising out of youruse of our data, or any suggestions for new topics or information that might lead to improvements, will help us toprovide a better service.

THE PREPARATION OF THIS DATA ITEM

The work on this particular Data Item which supersedes ESDU 66029, was monitored and guided by theAerodynamics Committee, which first met in 1942 and now has the following membership:

The technical work in the assessment of the available information and the construction and subsequent developmentof the Data Item method was undertaken by

ChairmanMr H.C. Garner Independent

Vice-ChairmanMr P.K. Jones British Aerospace plc, Aircraft Group, Manchester

MembersMr A. Condaminas Arospatiale, Toulouse, France.Mr E.A. Boyd Cranfield Institute of Technology Mr K. Burgin Southampton University Mr J.R.J. Dovey Independent Dr J.W. Flower Bristol UniversityMr A. Hipp British Aerospace plc, Dynamics Group, StevenageMr R. Jordan Aircraft Research Association Mr J. Kloos*

* Corresponding Member

Saab-Scania, Linkping, Sweden Mr J.R.C. Pedersen Independent Mr I.J. Rettie* Boeing Aerospace Company, Seattle, Wash., USAMr R. Sanderson Messerschmitt-Blkow-Blohm GmbH, Hamburg, Germany Mr A.E. Sewell* Northrop Corporation, Hawthorne, Calif., USAMr F.W. Stanhope Rolls-Royce plc, DerbyMr H. Vogel British Aerospace plc, Aircraft Group, Weybridge.

Mr R.W. Gilbey Senior Engineer.

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U. ESDU 86002DRAG AND PRESSURE RECOVERY CHARACTERISTICS OF AUXILIARY AIR INLETS AT SUBSONIC SPEEDS

CONTENTSPage

1. NOTATION AND UNITS (SEE SKETCH 1.1) 1

2. INTRODUCTION 5

3. QUALITATIVE CONSIDERATIONS OF DRAG CHARACTERISTICS 63.1 Basic Model 63.2 Pre-entry Forces 73.3 Drag Accounting 83.4 Spillage Drag 93.5 Comparison of Drag Characteristics 10

4. TWO-DIMENSIONAL BOUNDARY LAYER DATA 11

5. SCOOP INLETS 115.1 Drag 125.2 Pressure Recovery 145.3 Captured Mass Flow 15

6. FLUSH INLETS 166.1 Drag 176.2 Pressure Recovery 186.3 Comparison of Ramp Planform Effect 196.4 Captured Mass Flow 19

7. ACCURACY AND APPLICABILITY 227.1 Accuracy 227.2 Applicability 22

8. DERIVATION AND REFERENCES 248.1 Derivation 248.2 References 25

9. EXAMPLE 269.1 Scoop Inlet Designed for Full Mass Flow 26

9.1.1 Inlet size 269.1.2 Inlet drag 289.1.3 Pressure recovery 309.1.4 Summary 32i

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U. ESDU 860029.2 Flush Inlet Designed for Maximum Efficiency 33

9.2.1 Inlet size 339.2.2 Inlet drag 349.2.3 Pressure recovery 359.2.4 Summary 37

9.3 Flush Inlet Designed for Full Mass Flow 399.3.1 Inlet size 399.3.2 Inlet drag and pressure recovery 40

9.4 Comparison of Inlet Characteristics 42

FIGURES 1 to 20 43 to 62ii

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U. ESDU 86002DRAG AND PRESSURE RECOVERY CHARACTERISTICS OF AUXILIARY AIR INLETS AT SUBSONIC SPEEDS

1. NOTATION AND UNITS (see Sketch 1.1)

SI British

cross-sectional area of streamtube m2 ft2

forward projected area of diverter m2 ft2

external surface area of scoop inlet m2 ft2

forward projected area of scoop inlet outside highlight and excluding diverter, Am A1

m2 ft2

throat plane area m2 ft2

maximum outer cross-sectional area of inlet excluding diverter m2 ft2

inlet highlight capture area m2 ft2

drag coefficient,

drag coefficient at zero inlet mass flow, excluding skin friction and diverter drag

diverter drag coefficient

external skin friction drag coefficient of scoop inlet, CFAe /A1

spillage drag coefficient

incremental correction to drag coefficient for flush inlet

mean skin friction coefficient of flat plate based on surface area

flush inlet highlight height measured in inlet plane relative to ramp floor

m ft

scoop inlet lower highlight height relative to main surface m ft

distance between upper and lower highlights for scoop inlet, d1u d1l

m ft

scoop inlet upper highlight height relative to main surface m ft

A

Ad

Ae

Ap

At

Am

A1

CD D /0V02A1CD

CDd

CDF

CDsp

CDCF

d1 fl

d1 l

d1sc

d1u1

Issued April 1986With Amendments A to D, June 2004 - 62 pages

This page Amendment D

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U. ESDU 86002height of diverter measured from main surface to scoop lower surface

m ft

maximum height of scoop inlet relative to main surface m ft

maximum external height relative to ramp floor of flush inlet measured in inlet plane

m ft

distance between upper and lower internal surfaces in throat plane

m ft

drag N lbf

stream force, N lbf

height above flat plate, see Figures 1 to 4 m ft

correction factor allowing for effect of profile of fairing forebody on spillage drag of scoop inlet

correction factor allowing for effect of Mach number on for flush inlet

correction factor allowing for variation of with Mach number for scoop inlet

correction factor allowing for effect of mass flow on spillage drag

correction factor allowing for effect of ramp angle on flush inlet

factor applied to momentum flow to allow for ramp planform and lip shape of flush inlet

momentum flow correction factor

overall length of fairing of scoop inlet m ft

lip length, distance from entry plane to throat plane m ft

forebody length from lip to maximum height of scoop inlet m ft

length of ramp of flush inlet m ft

mass flow in streamtube entering inlet kg/s slug/s

mass flow in free-stream streamtube of cross-sectional area A1, kg/s slug/s

mass flow in free-stream streamtube of cross-sectional area , kg/s slug/s

dd

dmsc

dmfl

dt

D

F p p0( )A+h

kf

kMCD

kP PHt /PH0

ksp

k CD

k

K

lf

ll

lm

lr

m

m 0 0V0A1m 0( )t At0V0At2

Mach numberM


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U. ESDU 86002net propulsive force N lbf

static pressure N/m2 lbf/ft2

total pressure N/m2 lbf/ft2

total pressure in throat plane just aft of inlet lip N/m2 lbf/ft2

incremental correction to allow for variation of PHt with mass flow for scoop inlet

N/m2 lbf/ft2

correction factor allowing for effect of mass flow in calculation of

lip thickness at throat plane m ft

velocity of flow m/s ft/s

highlight width of inlet m ft

width of diverter m ft

coordinates of ramp planform m ft

ramp angle, see Sketch 1.1 degree degree

boundary layer thickness, where local velocity in boundary layer is 0.999V0

m ft

incremental correction to to allow for mass flow for flush inlet with curved-divergent ramp

incremental correction to to allow for inlet width for flush inlet with curved-divergent ramp

incremental correction to to allow for ramp angle for flush inlet with curved-divergent ramp

ram pressure efficiency, (PHt p0)/(PH0 p0)

maximum ram pressure efficiency of flush inlet with curved-divergent ramp

modified mass flow ratio for flush inlet with curved-divergent ramp,

value of modified mass flow ratio for flush inlet with curved-divergent ramp operating at maximum efficiency

boundary-layer momentum thickness m ft

density kg/m3 slug/ft3

forward component of pressure force on inside of streamtube N lbf

NPF

p

PH

PHt

PHt

rmf f l

t

V

w

wd

x ,y

mf

w

m

m d1 f l /m

0dt

m

3


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U. ESDU 86002rearward component of pressure force on outside of streamtube N lbf

momentum flow in streamtube entering inlet N lbf

momentum flow in free-stream streamtube of cross-sectional area A1 ,

N lbf

Subscripts

denotes free-stream conditions

denotes conditions far upstream on flat plate extending forward of inlet entry plane

denotes inlet entry plane

denotes inlet exit plane

denotes forward suction force

denotes choked flow

denotes fairing of scoop inlet

denotes flush inlet

denotes inlet running full, so that the free-stream boundary of the entering streamtube is unaffected by the inlet lip

denotes intrinsic net thrust

denotes pre-entry force

denotes ramp of flush inlet

denotes scoop inlet

denotes sharp lips

denotes spillage drag

denotes throat plane

denotes theoretical values in Figures 1 to 4

denotes theoretical values in Figures 1 to 4 with h = d1l

denotes theoretical values in Figures 1 to 4 with h = d1u

0

m 0V0

0

0 fp

1

9

c

ch

ff lfull

int

pre

r

sc

sh

sp

t

T

Tl

Tu4


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U. ESDU 86002

Sketch 1.1 Inlet geometries

2. INTRODUCTION

This Item provides data for the estimation of the drag and pressure recovery of small auxiliary air inletstotally or partially immersed in the boundary layer. The prediction method given is applicable at subsonicspeeds and is semi-empirical. In the main, it uses a theoretical calculation of two-dimensionalboundary-layer characteristics as a basis and modifies this in the light of the available experimental data.

A qualitative introduction to the drag characteristics of scoop and flush inlets is given in Section 3. Thedrag quantities to be estimated are identified and the variation of overall drag with mass flow ratio isconsidered.

Throat plane

Entry plane

Fairing

Diverter

Scoop inlet

Main surface

lflm

dm sc d1u d1 sc d1l

dd

dt

lrSee detail

ll

Flush inlet

dt

dm fl d1 fl

Throat plane

Entry plane

Main surface Main surface

Entry plane Throat plane

t

5

The theoretical data that are of use in the prediction of inlet performance are contained in Derivation 13which gives a two-dimensional analysis of the effect of boundary-layer thickness on the mass and

This page Amendment C

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U. ESDU 86002momentum flows of a given streamtube along a flat plate. Derivation 13 also gives the pressure recoveryjust aft of the inlet plane for a thin-lipped inlet running full. Numerical values of the theoretical data havebeen extracted from Derivation 13 and are reproduced herein, as described in Section 4.

The drag of a scoop inlet protruding from the main surface and having a rectangular, circular or semi-circularentry is considered in Section 5. The pressure recovery in the entry has been obtained from the theoreticalcalculations in Derivation 13 for full mass flow. A semi-empirical correction to allow for the loss in pressurerecovery at lower mass flows has been added from Derivation 19.

This Item provides a means of estimating the pressure recovery and drag characteristics of flush inlets withramp angles up to 11.5*. The drag of a flush inlet with a rectangular entry and an inlet ramp planform thatis rectangular or of the NACA curved-divergent type is considered in Section 6. A prediction of the pressurerecovery in the entry for a rectangular planform ramp has been obtained by making empirical correctionsto the data from Derivation 13. A separate method is given for predicting the pressure recovery for inletswith a NACA curved-divergent ramp planform.

For both scoop and flush inlets the experimental data available are very restricted in their range of flowconditions and inlet geometries, particularly in regard to lip profile, which is often sharp for theconfigurations tested. This must be remembered when considering the empirical corrections made to thebasic theoretical predictions. Section 7 on accuracy and applicability discusses this point in more detail.

Section 8 lists the Derivation and References that have been used in producing this Item.

Section 9 gives worked examples which compare the drag characteristics of scoop and flush inlets.

3. QUALITATIVE CONSIDERATIONS OF DRAG CHARACTERISTICS

3.1 Basic Model

This section makes a qualitative assessment along the lines set out in Derivation 18 in order to comparethe natures of the drag characteristics of scoop and flush inlets. Cross-sectional views of idealised,two-dimensional, scoop and flush inlets are shown in Sketch 3.1. The entry and exit planes are denoted (1)and (9), and station (0fp) is assumed to be far upstream on a flat plate extending forward of the inlet.

Sketch 3.1 Idealised inlets

* Item No. 03006 (Reference 32) extends the pressure recovery and drag prediction methods for rectangular planform flush inlets that turnthe flow through angles of up to 90 to the upstream flow direction.

prepre

c

f

r

(9)

(1)

Scoop inlet

Flush inletpre

pre(1)

(0fp)

(0fp)(9)6

This method was previously issued in Item No. 66029. That Item has now been withdrawn since it is superseded by the present Item. Adescription of the basic method in Item No. 66029 in its original form is available in Reference 27.


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U. ESDU 86002The streamtube passing through the inlet is bounded by the solid surfaces of the inlet and a free streamline,as shown dashed in the sketches for inlets running less than full. The net force on each inlet arises from thepressures acting internally and externally on the solid surfaces containing the streamtube.

The intrinsic net thrust of the inlets is equal to the difference in the stream force F that occurs betweenthe inlet entry and exit planes, so

(3.1)

where the stream force is given by

(3.2)

with constant velocity across the streamtube.

The net propulsive force, NPF , is equal to the difference between the intrinsic net thrust and the drag forcesacting either on the fairing of the scoop inlet, or on the ramp and lip of the flush inlet. Thus

, (3.3)

and

. (3.4)

For the scoop inlet, is the external pressure drag of the scoop fairing, including any lip or forebodysuction force (any skin friction drag is ignored at present). For the flush inlet, is the rearward forceassociated with flow over the entry ramp and represents the forward suction force on the lip.

3.2 Pre-entry Forces

It is convenient and customary to work with the net thrust relative to a plane far upstream (0fp), rather thanthe inlet entry plane (1). To achieve the necessary recasting of Equations (3.3) and (3.4), the change instream force between stations (0fp) and (1) is expressed in terms of forces acting forward of the entry plane.This brings in the concept of the equal and opposing pre-entry forces and that act on the freestreamline that defines the inlet streamtube. The force acts on the inside of the streamtube and ispositive upstream, and acts on the outside of the streamtube and is positive downstream. Therefore,

(3.5)

and

. (3.6)

Thus Equations (3.3) and (3.4) become

(3.7)

int

int F9 F1=

F p p0( )A+=

NPFsc int f F9 F1 f= =

NPFfl int r c F9 F1 r c= =

f r c

pre preprepreF1 F0fp( )sc pre pre= =

F1 F0fp( )fl pre r pre r= =

NPFsc F9 F0fp pre f=7


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U. ESDU 86002and

. (3.8)

Thus a pre-entry force is present for both types of inlet and this must be included in the drag calculation.

3.3 Drag Accounting

For the purposes of this Item it is assumed that all of the momentum of the air captured by the inlet is lostor is accounted for separately, so F9 can be set to zero.

(Derivation 16 gives an equation for the maximum possible thrust recovery in terms of the inlet ram pressureefficiency . This can be written

, (3.9)

but in a practical system internal pressure losses greatly reduce the actual thrust recovered.)

By Equation (3.2), , so that with F9 = 0

(3.10)

and

. (3.11)

If the momentum flow is taken as a ram drag forming part of the total inlet drag, then

(3.12)

and

. (3.13)

Because of the presence of a boundary layer the ram drag in the entering streamtube is less than it wouldbe if free-stream conditions prevailed upstream of the entry. If a factor K represents this loss, then

, (3.14)

where is the mass flow along the streamtube and V0 is the free-stream velocity.

NPFfl F9 F0fp pre c=

pre

F9 m V01 1 1 0.2M2+( )7/2 1[ ]+{ } 2/7

1 1 0.2M2+( ) 1------------------------------------------------------------------------------------

=

F0fp 0fp=

NPFsc 0fp pre f=

NPFfl 0fp pre c=

0fp

Dsc 0fp pre f+ +=

Dfl 0fp pre c+ +=

0fp K V0 m=

m8


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U. ESDU 860023.4 Spillage Drag

Scoop inlet

For the scoop inlet a forward suction force may develop on the lip or forepart of the fairing when the inletis not running full, see Item No. 81004 (Reference 28). This suction force can cause a total or partialcancellation of the pre-entry force . For inviscid flow over the fairing . For viscous flows,boundary layer growth reduces the suction force and causes a corresponding rise in the net force .As the mass flow is reduced from its full value the pressure gradients around the lip become more seriousuntil separation of the flow causes a very rapid decay of the suction force and a corresponding rise in

. The mass flow at which separation occurs depends on the lip geometry and Mach number, seefor example the data for free axisymmetric cowls in Item Nos 81024 and 84004 (Reference 29 andDerivation 24) which give information on this, together with theoretical values of in coefficient form.For ideal, sharp-lip inlets the flow is separated at all mass flows and there is no cancellation of the pre-entryforce.

If full is the pressure drag on the fairing at full mass flow, then the drag due to the flow that spills atlower mass flows can be defined using this as a datum value,

. (3.15)

Variations in Dsp with mass flow will then represent the large changes in .

Therefore, by Equations (3.12), (3.14) and (3.15)

. (3.16)

This can be expressed in coefficient form by dividing by or its equivalent , where is the free-stream mass flow through the inlet capture area A1, giving

. (3.17)

For an ideal, sharp-lip inlet CD sp will rise from zero for an inlet running full to a bluff-body value nearunity for zero mass flow. For a round-lip inlet the value will not rise until the mass flow has fallen to someintermediate value and will not rise to so high a value at zero mass flow. The variation of CDsp against massflow will change with Mach number. The datum coefficient CDf full is constant with mass flow.

Flush inlet

For the flush inlet, Derivation 18 considers an external streamtube that is defined by the streamline boundingthe flow entering the inlet, the aft surface of the inlet, and a streamline well above, as shown in Sketch 3.2.

pre f pre= pre f+

pre f+pre

f

Dsp pre f f full+=

pre f+

Dsc KV0m Dsp f full+ +=

0V20 A1 m 0V0 m 0

CDsc 2Km

m 0------ CDsp CDf full+ +=

AB

CA

C

B9

Sketch 3.2 Idealised flush inlet


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U. ESDU 86002If free-stream conditions exist at AA' and CC' and the flow between them is isentropic, then there is nochange in stream force and no net force on the streamtube. If the external streamline A'B'C' is well out inthe free stream and at ambient pressure there is no force on it, and consequently no force on the lowerboundary of the streamtube, ABC. Since the inlet after-surface is flush with the main surface and parallelto the free stream there is no force on the surface BC and so none on the surface AB.

Thus for the flush inlet will be small. For a round-lip inlet a suction force may also exist to oppose, as discussed above for scoop inlets. As there is no force on the inlet after-surface the net drag due to

spillage flow is which is zero for full mass flow. Using Equations (3.13) and (3.14) the flush inletdrag can then be written in coefficient form as

, (3.18)

where CDsp provides a small positive contribution at zero mass flow that decays as the mass flow increases.

3.5 Comparison of Drag Characteristics

Sketch 3.3 gives typical experimental drag characteristics from Derivation 16 for a sharp-lip scoop inletand a sharp-lip flush inlet of the same entrance area and a width to height ratio of 4. A tentative divisioninto component contributions, as suggested by Equations (3.17) and (3.18), is shown by dashed lines.

The scoop inlet has a much higher drag at low and intermediate mass flows, mainly due to its spillage dragcomponent. But for the flush inlet the ram drag contribution provides a reasonable firstapproximation to the total drag coefficient for all mass flows.

Sketch 3.3 Examples of typical drag characteristics of scoop and flush inlets

pre cpre pre c+

CDfl 2Km

m 0------ CDsp+=

2K m /m 0( )

0.00 0.25 0.50 0.75 1.000.0

0.5

1.0

1.5

2.0

Experiment

CD sp

CD fl

CD

Flush inlet

)2K( 0.m/ .m

.

m/ .m0

0.00 0.25 0.50 0.75 1.000.0

0.5

1.0

1.5

2.0

0.

m/ .m

CD sp

CD sc

Experiment

)2K( 0.m/ .m

CD f full

Scoop inlet10


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U. ESDU 860024. TWO-DIMENSIONAL BOUNDARY LAYER DATA

The two-dimensional theoretical analysis in Derivation 13 provides data that are required in the predictionof inlet drag and pressure recovery to account for the effect of boundary-layer thickness, , and velocityprofile on the flow properties of a streamtube of given height, h, passing along a flat plate. Of special interestfor the methods of this Item are the effects on mass and momentum flows. These properties can be expressedas a proportion of the values in streamtubes of equal thickness h with and without a boundary layer. For a1/7th power-law velocity profile, Figures 1 and 2 present and as functions of andM. The data in Figure 1 are of use in making an estimate of the size of an inlet required to capture a specifiedmass flow when running full. The data in Figure 2 are used in estimating the ram drag.

Information for calculating the total pressure recovery and the associated ram pressure efficiency for aninlet running full is also given in Derivation 13. Based on this, Figure 3 shows PHtT /PH0 and Figure 4shows as functions of and M for an inlet of height h.

Estimates of the boundary layer thickness may be made using Item No. 68020 (Derivation 20).

5. SCOOP INLETS

Section 5.1 describes a general method for predicting the drag of scoop inlets, which has been developedby using Equation (3.17) as a model. Experimental data for comparison are available for only a smallnumber of scoop geometries. Derivation 1 deals with inlets with a circular or a semicircular entry inlow-speed flow. Derivations 14 and 16 cover inlets with a high aspect-ratio, rectangular entry or a circular entry, in high-speed subsonic flow, . All of these inlets have sharp lips and noboundary-layer diverter. Sketches 5.1a to 5.1d illustrate typical geometries.

There is a tentative extension of the method, partly based on data from Item No. 84004 (Derivation 24), todeal with round-lip inlets for a range of fairing forebody shapes that include elliptical, NACA-1 series andlinear profiles. The comparative sketches in Item No. 84004 show that the first two profiles are fairly similar.A definition of the NACA-1 series profiles and a description of their performance advantages for isolated,axisymmetric cowls are given in Reference 25. For auxiliary inlets the simpler profiles may often be chosen.

Many of the correlating parameters in the method are presented separately for the particular cases ofrectangular inlets with , circular or semicircular inlets. To provide guidance for other entry shapessuch as rectangular inlets of low aspect-ratio a cross-plot against w2/A1 is also provided, but must be usedwith more caution than the individual data.

The method also covers the possible presence of a wedge-shaped boundary-layer diverter which, as an aidto pressure recovery, may be present to redirect all or the lower part of the boundary layer so that it is notcaptured by the inlet. The drag of the diverter is estimated as suggested in Derivation 21.

Section 5.2 describes the method for predicting the pressure recovery. This is based on theoretical data fromDerivation 13 and a semi-empirical correction provided by Derivation 19.

Section 5.3 explains how the size of the inlet may be determined once the mass flow to be captured hasbeen specified.

m T/m 0( ) T/0 /h

T /h

w2/A1 4( )0.55 M 0.9

w2/A1 411This page Amendment C

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U. ESDU 86002

Sketch 5.1 Example of inlet types tested

5.1 Drag

In Equation (3.17) the drag of a scoop inlet is divided into a ram drag, a spillage drag and a datum drag foran inlet running full. To these must be added the skin friction drag of the external surface of the scoop andthe drag of any diverter.

The ram drag term is obtained by replacing K in the first term of Equation (3.17) by

, (5.1)

where the theoretical momentum flow ratios and are obtained from Figure 2 withrespective heights h = d1u and d1l corresponding to the upper and lower highlights of the entry.Equation (5.1) is strictly valid only for rectangular inlets, but numerical checks have shown that it can beused directly for circular and semicircular inlets without any significant loss of accuracy.

The datum drag coefficient at full mass flow has been determined from the data in Derivations 1, 14 and 16.This is shown as CDf full in Figure 5. The data for circular, semicircular and rectangular inlets are indicatedindividually. For interpolation or extrapolation the data have been plotted against w2/A1. In all cases asmooth rear fairing is assumed with an overall fairing length . The experimental scatter shownby the broken curves is large because the effects of Mach number and the detailed design of the fairinghave not been eliminated. Data in Derivation 16 show an apparently random variation with Mach numberbetween the limits indicated. In the absence of additional data CDf full is assumed to apply to both sharp-lipand round-lip inlets.

The spillage drag term in Equation (3.17) is written as

CDsp = kf ksp sc ( CDf full)sh , (5.2)

and has been derived from experimental results in Derivations 1 and 16.

The quantity ( CDf full)sh is the difference between the drag coefficients at full and zero mass flowfor sharp-lip inlets and depends on Mach number. It is defined by the solid curves in Figure 6a for circular

a. Circular inlet

b. Semicircular inlet

c. Hooded semicircular inlet

d. Rectangular inlet

KscTu0---------

d1u Tl0--------

d1l 1

d1u d1l( )-------------------------=

Tu /0 Tl /0

lf 2.0dmsc

CD

CD12

and for semicircular or rectangular inlets, with broken lines indicating the experimental scatter. Thegeneralised trend with w2/A1 and M is shown in Figure 6b.


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U. ESDU 86002The factor ksp sc allows for the variation of the spillage drag with mass flow ratio. Figure 7a gives ksp sc forcircular inlets and Figure 7b gives ksp sc for semicircular or rectangular inlets. In both cases solid curvesdefine the factor and broken lines indicate the experimental scatter. The scatter includes the effect ofvariations with Mach number for , for which no systematic behaviour could be identified. Figure 7cgives the generalised trend of ksp sc with w

2/A1 and . In view of the small data base available fordefining ksp sc the precise shape of these curves is tentative but they show typical experimental behaviour.

The factor kf in Equation (5.2) models the significant reduction in spillage drag to be expected for inletswith a fairing forebody that has a smoothly shaped profile that allows the development of a forward suctionforce. The only experimental data for auxiliary inlets that permit this matter to be examined directly arethose in Derivation 1 for semicircular inlets with hooded forebodies and sharp lips, see Sketch 5.1c.Therefore, supplementary information at zero mass flow has been inferred from Item No. 84004 whichpredicts the spillage drag for free, axisymmetric cowls. Figure 8 gives values of kf that correspond to thereduction in spillage drag at zero mass flow for typical types of forebody. The factor depends on theparameters Ap/Am and lm/w, where Am is the maximum cross-sectional area of the inlet, Ap = Am A1, lmis the forebody length from the lip to the maximum height and w is the highlight width. Figures 8a and 8bdeal with elliptical and NACA-1 series forebody profiles for and , respectively, andFigure 8c deals with sharp-lip inlets where the forebody profile is linear. These data are a simpleapproximation to those in Item No. 84004 and they should be quite accurate for circular auxiliary inlets.The experimental data of Derivation 1 suggest that they also provide a good estimate for semicircular inlets.There are no data to verify the values of kf for rectangular or other inlet shapes but Figure 8 should give areasonable guide to the benefits of shaping the fairing forebody and rounding the inlet lips.

The skin friction drag of the scoop fairing is given by

CD F = CF Ae /A1 , (5.3)

where CF is the mean skin friction coefficient of a flat plate based on surface area, and may be obtainedfrom Item No. 68020 (Derivation 20) as a function of Reynolds number based on the fairing length lf .

Following the suggestion of Derivation 21 for two-dimensional wedge-shaped diverters it is assumed thatthe diverter drag coefficient is 0.25 based on its forward projected area, Ad , so that

CDd = 0.25 Ad /A1 . (5.4)

In some cases, for example circular inlets, the diverter may not be two-dimensional because of the lowersurface curvature of the fairing. For calculation purposes it is then sufficient to take Ad = wd dd .

The overall drag coefficient is thus

, (5.5)

where Ksc is from Equation (5.1),kf is from Figure 8,ksp sc is from Figure 7,( CDf full)sp is from Figure 6, CDf full is from Figure 5,

M 0.9m /m 0( )

M 0.6 M 0.2=

CDsc 2Kscm

m 0------ kf ksp sc CD CDf full( )sh CDf full CDF CDd+ + + +=

CD13

CDF is from Equation (5.3)


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U. ESDU 86002

5.2 Pressure Recovery

For a scoop inlet that is running full or nearly full, with no boundary-layer diverter, the pressure recoveryratio and the ram pressure efficiency in a throat plane just aft of the lip can be obtained from Figure 3 as

, (5.6)

and from Figure 4 as

, (5.7)

where in each case h = d1u , the height of the upper highlight of the entry.

If a diverter is present, so that the height of the entry lower highlight , the total pressure recovery atfull mass flow is obtained by using the data in Figure 3 and the equation

, (5.8)

where the theoretical pressure ratios PHt Tu /PH0 and PHt Tl /PH0 are obtained from Figure 3 with respectiveheights h = d1u and d1l .

For mass flows less than full there is an adverse interaction between the pre-entry pressure gradient andthe boundary layer. This distorts the boundary-layer profile and may cause separation. As a consequence,the pressure recovery in the throat is decreased by an amount . More importantly, the disturbednature of the entry flow reduces the efficiency of any diffuser downstream of the throat.

It is not possible to give a general quantitative prediction of these effects but for inlets without aboundary-layer diverter a limited amount of information is contained in Derivation 19. That Derivationreports a series of tests in which sharp-lip inlets, with d1l = 0, were mounted on a flat plate and the pressureloss due to the boundary-layer disturbance was measured at the end of diffusers of differing contours andlengths. Included in the test programme was a rectangular inlet followed by a straight, constant cross-sectionduct of length 4.5d1u. Derivation 19 contains a semi-empirical prediction for the pressure loss at thedownstream end of that duct and if it is assumed that the loss has the same form but a lower magnitude,then

, (5.9)

where is the boundary-layer momentum thickness, which can be estimated from Item No. 68020(Derivation 20), and kp is a function of Mach number given in Figure 9.

and CDd is from Equation (5.4).

PHtPH0----------

sc full

PHt TuPH0

----------------=

sc full Tu=

d1l 0

PHtPH0----------

sc full

PHt TuPH0

---------------- d1u

PHt TlPH0

-------------- d1l

1d1u d1l( )

-------------------------=

PHt

PHt

PHtPH0------------

kp

d1u-------- 1 m

m 0------

3

14This page Amendment C

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U. ESDU 86002If a boundary-layer diverter is present the interaction between the pressure gradient and the boundary layerbecomes unimportant and

. (5.10)

The total pressure recovery ratio is

. (5.11)

This is converted to a ram pressure efficiency through the equation

. (5.12)

The data available to verify these equations from pressure measurements made in the inlet throat are verylimited. In Derivation 18 a few data are collected for rectangular inlets, without diverters, operating at fullmass flow. Those data are in good agreement with the predictions of Equation (5.6).

5.3 Captured Mass Flow

The mass flow ratios in Figure 1 allow the effect of boundary-layer thickness to be considered whenestimating the size of an inlet that at full mass flow captures air at a specified mass flow.

For a scoop inlet running full

, (5.13)

where and are calculated for h = d1u and d1l respectively. Although strictly valid onlyfor rectangular inlets numerical checks have shown that Equation (5.13) can be used for circular andsemicircular inlets without any significant loss of accuracy.

Equation (5.13) is used to calculate d1u once and the free-stream conditions are known, thecross-sectional shape of the entry specified, and d1l chosen, say, as a fixed proportion of the boundary-layerthickness. This is accomplished by first setting and obtaining a first estimate of d1u in theimplied absence of a boundary layer. Equation (5.13) can then be used to establish the shortfall in capturedmass flow that will be occasioned by the boundary layer. The necessary increase in A1 and hence d1u , thatis needed to offset this can then be estimated. This procedure is demonstrated in Section 9.1.1 of the examplefor a circular scoop.

Assuming the cross-sectional area of the inlet is least in the throat plane just aft of the lip, Derivation 22with Reference 28 gives the theoretical mass flow for a choked inlet,

. (5.14)

PHtPH0------------ 0=

PHtPH0----------

sc

PHtPH0----------

sc full

PHtPH0------------+=

sc1 0.2M2+( )7/2 PHt /PHo( )sc 1

1 0.2M2+( )7/2 1---------------------------------------------------------------------------=

m 0V0A1m Tum 0---------

d1u m Tlm 0--------

d1l 1

d1u d1l( )-------------------------=

m Tu /m

0 m

Tl /m

0

m

A1 m /0V0=

m ch-------------

m ch-------------------------

0.579-------------

PHt----------

1 0.2M2+( )3==15

m 0( )t At/A1( )m 0 M PH0

sc


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U. ESDU 860026. FLUSH INLETS

Section 6.1 describes how a general method for predicting the drag of certain types of flush inlets has beendeveloped by taking Equation (3.18) as an initial model. Empirical corrections are based on the availableexperimental data, but it has only been possible to produce a method for a limited range of inlet geometries.The method in Section 6.1 is applicable for ramp angles . Item No. 03006 (Reference 32) maybe used to predict the drag of rectangular planform flush inlets with geometries that turn the flow throughlarger angles (up to 90 to the upstream flow direction).

The two basic types of inlet considered are shown in Sketch 6.1. The first has a ramp of rectangular planform,which has the advantage of ease of manufacture. The second has a NACA curved-divergent ramp with ahighly swept planform that causes strong vortices to develop along the ramp sides. These thin the boundarylayer in the centre of the ramp and so reduce the pressure loss due to the boundary layer. Although thereare additional losses due to the vortices themselves, the net result is a pressure recovery superior to that ofother flush inlets and which exceeds the theoretical two-dimensional pressure recovery. See Derivations 2to 9, 11, 12, 17 and 23 and References 26 and 30. For best pressure recovery an inlet width to depth ratioof about 4 and a ramp angle of 7 are desirable.

Equation (3.18) has been used to analyse the two sets of data that are provided by the parametric testprogrammes reported in Derivations 3 and 16. Derivation 3 treats inlets of aspect ratio 2, 4 or 6 with roundedlips in low-speed flow with and , and Derivation 16 treats inlets of aspect ratio 4 withsharp lips in high-speed subsonic flows, . No other collections of systematic data are

Ramp coordinates for NACA curved-divergent planformx/lr 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

2y/w 1.0 0.996 0.916 0.766 0.614 0.466 0.388 0.312 0.236 0.158 0.085

Sketch 6.1 Flush inlets

11.5

A

B wlr

B

AA

A

B B

Section B-B

Rectangular planform

yx

Plan views

Section A-A

w

ll /t 1 dt /t 20.55 M 0.9 16

available, although there are a few isolated data from particular tests.


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U. ESDU 86002Section 6.2 describes how the pressure recovery is calculated for the two types of flush inlet. For therectangular ramp inlet pressure recovery is obtained to a first approximation. The method uses thetheoretical predictions of Derivation 13 together with an empirical correction to allow for the variation withmass flow based on data from Derivation 3. For the curved-divergent ramp inlet a separate method is given(see footnote in Section 2).

Section 6.3 presents a comparison of the drag and pressure recovery characteristics of inlets with differingramp planforms.

Section 6.4 describes how the size of a flush inlet is designed to accommodate a specified mass flow. Aspecial method is given for inlets with a curved-divergent ramp operating at maximum efficiency.

6.1 Drag

Following Equation (3.18) the drag of a flush inlet is divided into a ram drag term and a spillage drag term,the former dominating at moderate to high mass flows.

The ram drag term is obtained by replacing K in the first term in Equation (3.18) by

, (6.1)

where is the theoretical momentum flow ratio from Figure 2 for a height h = d1 fl , correspondingto the inlet highlight height.

The factor is an empirical constant that has been deduced from the experimental data for inlets runningfull or near to their maximum mass flow. For inlets with round lips, and regardless of ramp profile, the datain Derivation 3 suggest that for low-speed flow and ramp angles

, (6.2)

but there is evidence that decreases by about 10 per cent for . The data in Derivation 16suggest that for sharp-lip inlets with rectangular ramps, with , Equation (6.2) is valid for .Figure 10a summarises the general behaviour of that has been inferred by combining these data. Forsharp lip inlets with a curved-divergent ramp, however, data in Derivation 16 show that increases withMach number from a value of 0.8 at M = 0.55 to a value of 1.0 at M = 0.9, as shown in Figure 10b. Forhigh ramp angles a reduction in similar to that shown in Figure 10a has been applied. Althoughdetermined at high mass flows, these values of should provide a good approximation to theboundary-layer loss for all mass flows.

The spillage drag coefficient has been determined empirically from data in Derivations 3 and 16 bycorrelating a drag coefficient, , at zero mass flow and low Mach number in terms of lip shape for a 7ramp, as shown in Figure 11. Because there is only one set of experimental data available for inlets withround lips, the parameter defining the lip shape, (dmfl d1fl)/ll, is simple and carries the assumption thatthe profile is approximately elliptical. An interpolation between the experimental results for round-lip andsharp-lip inlets and an extrapolation to circular-profile lips are shown dashed. Two factors, and kM givenin Figures 12 and 13, are applied to , to allow for ramp angles and Mach numbers .The factors and kM have been derived solely from data on round-lip and sharp-lip inlets, respectively.Moreover, only data for curved-divergent ramp planforms are available for defining .

Kfl kT0-------

=

T /0

k 9

k 1=

k 11.5= 7= M 0.9k

kk

k

CDfl

kCDfl 7 M 0.55

kk17


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U. ESDU 86002But in the absence of more widespread information these two factors are assumed to be of generalapplication. A third empirical factor ksp fl, given in Figure 14, allows for the decay in spillage drag thatoccurs as the mass flow increases. Once again, the data for constructing this figure are very restricted inthe range of lip shapes tested. The curve for M = 0.2 has been obtained for round lips and a rectangularplanform ramp. Those at higher Mach numbers have been obtained for sharp lips for both rectangular andcurved-divergent ramp planforms. The precise shape of the curves and the points at which ksp fl becomeszero must be regarded as tentative, but they are typical experimental results. Because the spillage dragcomponent is small any errors due to this simple approach should be tolerable, especially since thepredictions for zero mass flow are quite well defined.

Comparison of the total predicted ram and spillage drags with experimental data shows that at high massflows the inlet drag is sometimes underpredicted by an amount that increases with Mach number. To allowfor this a correction is introduced, as given in Figures 15a and 15b for rectangular andcurved-divergent ramps respectively. Here, again, the curves at M = 0.2 are derived from one test series oninlets with round lips and those for higher Mach numbers from inlets with sharp lips, so the curves provideno more than an approximate correlation.

The drag coefficient of the flush inlet is then

, (6.3)

Item No. 03006 (Reference 32) may be used to obtain the drag coefficient of rectangular planform inletsthat turn the flow through angles exceeding 11.5, up to a maximum of 90 to the upstream flow direction.

6.2 Pressure Recovery

The ram pressure efficiency for an inlet with a rectangular planform ramp and which is running full can,to a first approximation, be obtained as from the theoretical results in Figure 4, where h = d1 fl is thehighlight height of the inlet. The experimental, low-speed data from Derivation 3 for round-lip inlets havebeen used to produce a factor rmf that allows for the reduction in efficiency that occurs as the mass flowdecreases. This is given in Figure 16.

The ram pressure efficiency is calculated as

, (6.4)

where corresponds to h = d1fl. These data are also assumed to apply to sharp-lip inlets.

Item No. 03006 (Reference 32) may be used to predict the ram pressure efficiency of more complexrectangular flush inlets geometries, which turn the flow through angles up to a maximum of 90 to theupstream flow direction and includes an effect on Mach number that is not present in the first approximation

where Kfl is from Equation (6.1), is from Figure 12,

kM is from Figure 13, ksp fl is from Figure 14,

is from Figure 11,and is from Figure 15.

CD

CDfl 2Kflm

m 0------ kkMksp fl CDfl CD+ +=

k

CD flCD

T

fl rmf T=

T18

that is provided in Equation (6.4).


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U. ESDU 86002The ram pressure efficiency of inlets with a curved-divergent ramp has been obtained by analysing the datain Derivations 2 to 9, 11, 12, 17 and 23 (see footnote in Section 2). In those Derivations pressuremeasurements made in the throat plane are given for inlets with lip profiles that are approximately ellipticalwith and . The experimental data show that for the intake performance isindependent of Mach number. Derivation 22 suggests that the efficiency only decreases by about 1 per centfrom M = 0.8 to M = 0.9, which is supported by experimental results in Derivations 7 to 10.

For inlets with a curved-divergent ramp the value of is principally determined by the boundary-layermomentum thickness to throat height ratio , where may be estimated from Item No. 68020(Derivation 20). For fixed geometry and for a given value of , varies with mass flow and has amaximum value . There is a corresponding value, , of the modified mass flow ratio when . The solid curve in Figure 17 shows the influence of on for , w/dt = 4 and

. Dashed lines indicate a per cent scatter band that contains 80 per cent of the availableexperimental data. Similarly, Figure 18 gives as a function of with a per cent scatter bandthat contains 60 per cent of the experimental data.

Small incremental changes are employed to allow for the effect of parameters other than and thegeneral expression for is

, (6.5)

where is a correction in Figure 19a dependent on the change in modified mass flow ratio, .The last two terms, from Figures 19b and 19c respectively, allow for ramp angles and inlet aspect ratiosother than the optimal values and w/dt = 4.

For sharp-lip inlets there are no experimental data on the pressure recovery in the throat or entry plane, allavailable measurements having been made well downstream. However, in the absence of evidence to thecontrary it is reasonable to assume that the curved-divergent ramp planform will provide the same benefitover a rectangular ramp regardless of lip profile. Figures 17 to 19 can therefore be used for both sharp-lipand round-lip inlets.

6.3 Comparison of Ramp Planform Effect

An example of the experimental values of drag coefficient and ram pressure efficiency for different rampplanforms is given in Sketch 6.2. These data, taken from Derivation 3 in low-speed flow, compare arectangular, a curved-divergent and the intermediate case of a straight-divergent ramp with a closure ratioof 0.25. The straight-divergent ramp gives a slightly smaller drag than the curved-divergent ramp and isnearly as efficient in terms of pressure recovery except at low mass flows. There are insufficient data todevelop a separate prediction method for a straight-divergent ramp, but Sketch 6.2 may be used as a guideto modify the results predicted for a curved-divergent ramp.

6.4 Captured Mass Flow

For a flush inlet with a rectangular ramp that is operating at full mass flow, the captured mass flow can bewritten

, (6.6)

ll /t 1 dt /t 2 M 0.8

fl/dt /dt flm m m d1fl /m 0dt=fl m= /dt m 7=M 0.8 5

m /dt 10

/dtflfl m mf w+ + +=

mf m 7=

m 0V0A1m Tm 0-------

0V0wd1fl

m Tm 0-------

= =19


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U. ESDU 86002where the mass flow ratio is obtained from Figure 1 with h = d1fl . With the free-stream conditionsknown and w/d1fl chosen, successive numerical calculations using Equation (6.6) can be made to determinethe value of d1fl necessary to accommodate a specified mass flow.

A flush inlet with a curved-divergent ramp that is operating at full mass flow may be sized in a similarmanner, although because the ramp planform mitigates the effect of the boundary layer there may be someoverestimate of the inlet dimensions. The procedure is illustrated in Section 9.3.1 of the example.

For a flush inlet with a curved-divergent ramp that is operating at maximum efficiency there is a specialcorrelation for determining the inlet size. Figure 20 plots against the mass flow parameter when . If and are known this defines dt without the necessity of repetitive numericalcalculations. Although developed primarily for inlets with w/dt = 4, for which the data are an alternativepresentation of those in Figure 18, Figure 20 includes correlations for w/dt = 2 and 6 although these arebased on far fewer data points and extrapolated trends are shown as dashed lines. With w/dt specified andthe lip geometry known, A1 can be calculated. This is demonstrated in Section 9.2.1 of the example. It isassumed that a sharp-lip inlet can be sized in the same manner.

If the cross-sectional area of the inlet is least in the throat plane just aft of the lip the theoretical mass flowratio at which the inlet is choked is

, (6.7)

which is Equation (5.14) rewritten in terms of ram pressure efficiency. It is pointed out in Derivation 22that the mass flow is more likely to be restricted by the internal pressure than by choking.

m T/m 0

/dt m /0V02fl m= m

m chm 0( )t

-------------

m chAt/A1( )m 0

-------------------------

0.579M

------------- 1 0.2M2+( )3 fl1 fl( )

1 0.2M2+( )7/2-----------------------------------+

==20


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U. ESDU 86002

.

m

m0

.

C

BA

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

CD f l

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

1.8

2.0

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

0.0

0.2

0.4

0.6

0.8

1.0

m

m0.

.

00

00

C

B

A

C

B

A

w0.083w0.25w

ABC21

Sketch 6.2 Comparison of ramp planform effects


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U. ESDU 860027. ACCURACY AND APPLICABILITY

7.1 Accuracy

For the experimental data that have been studied the methods developed in this Item predict the drag ofauxiliary scoop inlets to within about per cent at zero mass flow, the accuracy improving to about

per cent at full mass flow, For flush inlets the corresponding figures are and per cent. Asmentioned in previous sections some caution is necessary because many of the test data are for sharp-lipinlets. In practice it is expected that there will be some rounding of the lip similar to that employed onaxisymmetric engine cowls. For scoop inlets it has not been possible to examine the effect of the aft fairingin detail and a smooth shape must be assumed.

Few data are available for assessing the accuracy with which the pressure recovery can be predicted forscoop inlets or flush inlets with a rectangular ramp or sharp lips. For scoop inlets a few data in Derivation 18suggest that for full mass flow is predicted to within about per cent. For flush inlets with rectangularramps and round lips the experimental data are predicted to within about per cent. For flush inlets withcurved-divergent ramps and round lips data are more abundant and comparisons with many test resultsshow an overall accuracy of per cent.

7.2 Applicability

The ranges of inlet geometries and flow conditions for the systematic sets of test data that have been usedto develop the methods in this Item are set out in Tables 7.1 and 7.2. For scoop inlets the method has beenextended from the basic data so as to include the effects of boundary-layer diverters and round lips. Althoughthe approach adopted is thought to be reasonable some caution is necessary in the absence of directcomparisons of prediction and experiment for such inlets.

All of the systematic data have been taken from tests on inlets mounted on flat plates parallel to a uniformfree-stream. In practice an inlet may be mounted on a main surface that is inclined to the free stream. Themethod is applicable for boundary layers upstream of the inlet exposed to neutral or favourable pressuregradients. It is evident from Reference 31 that with a favourable pressure gradient the boundary layer maybe approximated by a one-seventh power law velocity profile. However the velocity profile diverges froma one-seventh power law profile when the boundary layer is exposed to an adverse pressure gradient. It isassumed that the inlet is aligned with the local flow so there is no yaw angle.

The rearward force due to the inlet can be resolved into the drag direction to account for the geometriceffect on the inclination of the main surface.

To allow for the pressure field the local Mach number, density and velocity at the edge of the boundarylayer should be substituted for M, and V0 when estimating drag and pressure recovery. It must howeverbe noted that, as discussed in Reference 30, the use of local flow conditions will only provide an acceptableapproximation to drag calculations for small pressure gradients, because there is a magnification ofexcrescence drag due to the behaviour of the boundary layer behind an excrescence and this can be largefor severe pressure gradients.

If the inlet is yawed with respect to the local flow a decrease in pressure recovery can be expected. This isdifficult to quantify but, as an example, Derivation 12 indicates that a yaw angle of 5 reduces by4 per cent for a flush inlet with a curved-divergent ramp and there is a corresponding increase ofapproximately 10 per cent in the value of mass flow required for maximum efficiency.

1510 15 5

53

5

0

m22


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U. ESDU 86002TABLE 7.1 Ranges of test parameters for drag measurements

SCOOP INLETS

Derivation Lip shape Entry cross-section lf /dmsc M

1

sharpsharp

circularsemicircular(including hoodedforebodies )

2 to 4.52 to 4.5

14 sharp rectangular 5 to 10 0.64

16 sharpsharp

circularrectangular (w/d

m sc = 4)

510

0.55 to 0.90.55 to 0.9

FLUSH INLETS

Derivation Lip shape Entry cross-section w/d1fl M

3 roundround

rectangularcurved-divergent

2, 4, 64

75 , 7 , 9 , 11.5

16 sharpsharp

rectangularcurved-divergent

44

77

0.55 to 0.90.55 to 0.9

TABLE 7.2 Ranges of test parameters for pressure recovery measurements

FLUSH INLETS

Derivation Lip shape Ramp profile w/dt M

3 round curved-divergent 41 to 65 to 11.5

70.05

0.025 to 0.135

4 round curved-divergent 4 7 0.02 to 0.03 0.2

7 round curved-divergent 4 6.5 0.012 to 0.044 0.8 to 0.9

8 round curved-divergent 4 7 0.02 to 0.065 0.3 to 0.9

9 round curved-divergent 4 7 0.03 0.8 to 0.9

11, 12, 17 round curved-divergent 3.8, 4, 4.6 7, 8 0.04 to 0.424 0.1

/d1u

0 Ap/A1 0.5

0.40.4 to 0.7 0.2 0.2

w/dmsc 5 to 10( ) 0.6 to 0.70.40.8

/d1fl0.4 to 0.80.5

0.230.23

0.80.8

/dt0.230.2323


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U. ESDU 860028. DERIVATION AND REFERENCES

8.1 Derivation

The Derivation lists selected sources that have assisted in the preparation of this Item.

1. ROGALLO, F.M. Internal flow systems for aircraft, NACA Rep. 713, 1941.

2. FRICK, C.W. DAVIS, W.F. RANDALL, L.H. MOSSMAN, E.A.

An experimental investigation of NACA submerged duct entrances.NACA RM 5I20 (TIL 1778), 1945.

3. MOSSMAN, E.A.RANDALL, L.H.

An experimental investigation of the design variables for NACAsubmerged duct entrances. NACA RM A7I30 (TIL 1563), 1947.

4. DELANY, N.K. An investigation of submerged air inlets on a -scale model of a typicalfighter-type airplane. NACA RM A8A20 (TIL 1822), 1948.

5. MARTIN, N.J. HOLZHAUSER, C.A.

An experimental investigation at large scale of several configurations ofan NACA submerged air intake. NACA RM A8F21 (TIL 1937), 1948.

6. HALL, C.F. FRANK, J.L.

Ram-recovery characteristics of NACA submerged inlets at highsubsonic speeds. NACA RM A8I29 (TIL 1970), 1948.

7. AXELSON, J.A. TAYLOR, R.A.

Preliminary investigation of the transonic characteristics of an NACAsubmerged inlet. NACA RM A50C13 (TIL 2393), 1950.

8. FRANK, J.L. Pressure distribution and ram-recovery characteristics of NACAsubmerged inlets at high subsonic speed. NACA RM A50E02 (TIL2412), 1950.

9. SELNA, J. SCHLAFF, B.A.

An investigation of the drag and pressure recovery of a submerged inletand a nose inlet in the transonic flight range with free-fall models.NACA RM A51H20 (TIL 2944), 1951.

10. FRANK, J.L. TAYLOR, R.A.

Comparison of drag, pressure recovery, and surface pressure of a scooptype inlet and a NACA submerged inlet at transonic speeds. NACA RMA51H20a (TIL 2945), 1951.

11. The effect of boundary layer thickness on the ram-recovery of flushintakes. A.V. Roe and Co. Ltd. Wind Tunnel Report No. 698/65, 1952.

12. The effect of boundary layer thickness on the ram recovery ratio of anNACA flush intake. A.V. Roe and Co. Ltd, Wind Tunnel ReportNo.698/80, 1953.

13. SIMON, P.C. KOWALSKI, K.L.

Charts of the boundary-layer mass flow momentum for inletperformance analysis, Mach number range, 0.2 to 5.0. NACA tech.Note 3583, 1955.

14. HEARTH, D.P. CUBBISON, R.W.

Investigation at supersonic and subsonic Mach numbers of auxiliaryinlets supplying secondary air flow to ejector exhaust nozzles. NACARM E55J12a (TIL 4942), 1956.24


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U. ESDU 86002

8.2 References

The References list selected sources of information supplementary to that given in this Item.

15. DENNARD, J.S. A transonic investigation of the mass-flow and pressure recoverycharacteristics of several types of auxiliary air inlets. NACA RML57B07 (TIL 5490), 1957.

16. DENNARD, J.S. The total pressure recovery and drag characteristics of several auxiliaryinlets at transonic speeds. NACA Memo 12-21-58L (TIL 6469), 1958.

17. Tests on the NACA-type flush air intake fitted to the SundstrandAlternator and oil-cooler units. A.V. Roe and Co. Ltd, Wind Tunnelreport ARD/WT/698/179, 1958.

18. LEYLAND, D.C. The drag of scoop and flush auxiliary air inlets. BAC (Preston) RepP.M. 188, 1966.

19. SEDDON, J. Boundary layer interaction effects in intakes with particular reference tothose designed for dual subsonic and supersonic performance. RAEtech. Rep. 66099, 1966.

20. ESDU The compressible two-dimensional turbulent boundary layer, both withand without heat transfer, on a smooth flat plate, with application towedges, cylinders and cones. Item No. 68020. Engineering SciencesData Unit, London, 1968.

21. DAVENPORT, C. A further investigation of the drag at subsonic speeds of side intakeboundary layer diverters. S & T Memo 7/68, Ministry of Technology1968.

22. BEST, M.S. NACA submerged air inlets. BAC (Filton) Aero TN/MSB/LJF/213, c.1969.

23. ARA Unpublished information from Aircraft Research Association Ltd,Bedford.

24. ESDU Estimation of spillage drag for a wide range of axisymmetric intakes at. Item No. 84004. Engineering Sciences Data Unit, London,

1984.

25. BAALS, D.D. SMITH, N.F.WRIGHT, J.B.

Development and application of high-critical-speed nose inlets. NACARep. 920, 1948.

26. SACKS, A.H. SPREITER, J.R.

Theoretical investigation of submerged inlets at low speeds. NACAtech. Note 2323, 1951.

27. McCREATH, K.M. WARD SMITH, A.J.

Boundary layer influence on the performance of submerged intakes. J1.R.Ae.S., 1967.

28. ESDU Mass flow and momentum functions for one-dimensional flow of gas inducts. Item No. 81004. Engineering Sciences Data Unit, London, 1981.

M 125


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9. EXAMPLE

Compare the drag and ram pressure efficiency of

(i) a circular scoop inlet designed to operate at full mass flow,

(ii) a flush inlet with a curved-divergent ramp designed to operate at maximum ram pressure efficiency,

and

(iii) a flush inlet with a curved-divergent ramp designed to operate at full mass flow.

All three inlets are aligned to the free stream and designed to capture air at a given flow of 0.0160 slug/sat their operating points. The local flow conditions outside the boundary layer are close to the free-streamvalues M = 0.8, slug/ft3 , V0 = 995 ft/s and a Reynolds number per unit length of7.5 106 ft1. The boundary-layer thickness is 0.1 ft and the boundary-layer momentum thickness is 0.01 ft.

The overall length of the circular scoop inlet is lf = 0.6 ft and its maximum height is dmsc = 0.19 ft. Thefairing forebody, of length lm = 0.06 ft, has a semi-elliptical external profile. The internal lip has aquarter-circle profile of raduis 0.002 ft between the lip highlight and the throat plane, which is 0.002 ft aftof the entry plane. A boundary-layer diverter, situated 0.006 ft aft of the inlet entry, raises the lower liphighlight above the main surface so that d1l = 0.02 ft.

Both of the flush inlets have a ramp angle of and a throat aspect ratio w/dt = 4. In each case the lipis assumed to be elliptically rounded with the lip length and thickness in the throat plane both equal to aquarter of the throat diameter (ll = t = 0.25 dt).

9.1 Scoop Inlet Designed for Full Mass Flow

9.1.1 Inlet size

A first approximation to the magnitude of the inlet capture area is obtained by ignoring the boundary layerand setting

ft2 .

29. ESDU Drag of axisymmetric cowls at zero incidence for subsonic Machnumbers. Item No. 81024. Engineering Sciences Data Unit, London,1981.

30. AGARD Aircraft excrescence drag. AGARDograph No. 264, 1981.

31. CASTILLO, L.GEORGE, W.K.

Similarity analysis for turbulent boundary layer with pressure gradient:outer flow. AIAA Journal, Vol. 39, No. 1, pp 41-47, January 2001.

32. ESDU Subsonic drag and pressure recovery of rectangular planform flushauxilary inlets with ducts at angles up to 90 degrees. Item No. 03006.Engineering Sciences Data Unit, London, 2003.

0 0.890 10 3=

7=

A1m

0V0------------

0.01600.890 10 3 995-------------------------------------------- 0.0181= = =26


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U. ESDU 86002For a circular scoop,

ft.

With d1l = 0.02 ft it follows that d1u = 0.172 ft.

The effect of the boundary layer on the captured mass flow can be examined by means of Equation (5.13)

.

The two mass flow ratios and are determined from Figure 1 for the values and , respectively, with M = 0.8.

Equation (5.13) then gives

slug/s.

Thus the boundary layer causes a shortfall of 5 per cent below the desired capture flow of 0.0160 slug/s. IfA1 is increased by 5 per cent,

ft2 ,

giving ft,

and ft,

with the throat diameter ft,

giving ft2.

A check using Equation (5.13) with the increased dimensions confirms that the desired mass flow isachieved with

slug/s,

and this corresponds to a mass flow ratio

.

d1sc d1u d1l 2 0.0181/( ) 0.152= = =

m 0V0A1m Tum 0---------

d1um Tlm 0--------

d1l 1

d1u d1l( )-------------------------=

m Tu /m 0 m Tl /m 0/h /d1u 0.1/0.172 0.581= = = /h /d1l 0.1/0.02 5.0= = =

m 0.890 10 3 995 0.0181 0.915 0.172 0.655 0.02{ } 10.172 0.02( )---------------------------------= = 0.0152

A1 0.0181 1.05 0.0190= =d1sc 0.156=

d1u 0.176=

dt 0.156 2 0.002 0.152==At dt2/4 0.1522/4 0.0181== =

m 0.890 10 3 995 0.0190 0.917 0.176 0.655 0.02{ } 10.176 0.02( )---------------------------------= 0.0160=

m

m 0------

0.01600.890 10 3 995 0.0190------------------------------------------------------------------ 0.950= =27This page Amendment D

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U. ESDU 860029.1.2 Inlet drag

The overall drag coefficient is given by Equation (5.5)

.

(a) For the ram drag component, the first term on the right-hand side, the value of Ksc is obtained fromEquation (5.1)

.

The two momentum flow ratios and are obtained from Figure 2 for the values and , respectively, with M = 0.8. Thus Equation (5.1)

becomes

.

The ram drag component of Equation (5.5) can then be summarised as follows.

(b) For the spillage drag coefficient, the second term, the value of kf allows for the effect of forebodyprofile in terms of Mach number and the geometric parameters lm /w and Ap /Am.

For the present circular scoop, w = d1sc = 0.156 ft and

.

The maximum diameter of the scoop is d1sc + 2(dmsc d1u), so that

0 0.2 0.4 0.6 0.8 1.0

2ksc 0 0.366 0.731 1.097 1.462 1.828

.

CDsc 2Kscm

m 0------ kf ksp sc CD CDf full( )sh CDf full CDF CDd+ + + +=

KscTu0---------

d1uTl0--------

d1l 1

d1u d1l( )-------------------------=

Tu /0 Tl /0/h 0.1/0.176 0.568= = /h 0.1/0.02 0.5= =

Ksc 0.863 0.176 0.462 0.02{ } 10.176 0.02( )---------------------------------= 0.914=

m /m 0m /m 0( )

lmw-----

0.0600.156------------- 0.385= =

ApAm------- 1

A1Am-------

= 1

d21scd1sc 2 dmsc d1u( )+[ ]2

-------------------------------------------------------=

1 0.1562

0.156 2 0.190 0.176( )+[ ]2------------------------------------------------------------------=

0.281=28This page Amendment C

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U. ESDU 86002Then from Figure 8a for an elliptical profile with M = 0.8, cross-plotting against lm /w gives

kf = 0.275.

The factor ksp sc allows for the variation of spillage drag with mass flow ratio. For a circular inlet Figure 7agives the following.

The value of the spillage drag at zero mass flow for sharp lip inlets is obtained from Figure 6a at M = 0.8,

( CDf full)sh = 1.12.

The spillage drag coefficient as a function of mass flow can then be calculated.

(c) The datum drag coefficient at full mass flow, the third term, is obtained from Figure 5 for a circularinlet as

CDf full = 0.300.

(d) To estimate the skin friction drag coefficient, the fourth term, the external surface area of the scoop,Ae, and the mean skin friction coefficient CF are required. For the present example Ae is taken to be 0.3 ft

2.

With lf = 0.6 ft the Reynolds number based on the fairing length is 7.5 106 0.6 = 4.50 106. For this

value and M = 0.8 Item No. 68020 gives CF = 3.3 103

, whence

.

(e) The diverter drag coefficient, the fifth and last term, is obtained from Equation (5.4)

.

The front of the diverter is situated 0.006 ft aft of the inlet entry, and the elliptical profile of the lower lipdetermines its height as dd = 0.014 ft. If the diverter width is equal to the maximum external diameter of

0 0.2 0.4 0.6 0.8 1.0

ksp sc 1.0 0.625 0.360 0.165 0.040 0

0 0.2 0.4 0.6 0.8 1.0

kf ksp sc ( CD f full)sh 0.308 0.193 0.111 0.051 0.012 0

m /m 0

CD

m /m 0CD

CDF CF AeA1------=

3.3 10 3 0.30.0190

------------------------------------- 0.052= =

CDd 0.25AdA1------=29This page Amendment C

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U. ESDU 86002the inlet wd = w + 2(dmsc d1u) = 0.184 ft, Ad can be approximated as

ft2 .

Therefore .

(f) The total drag coefficient as a function of mass flow ratio is obtained by summing the componentsestimated above, as set out in Table 9.1.

9.1.3 Pressure recovery

The total pressure recovery for the inlet running full is given by Equation (5.8)

,

where the two pressure ratios PHt Tu /PH0 and PHt Tl /PH0 are determined from Figure 3 for the values and , respectively, with M = 0.8.

Equation (5.8) then becomes

.

Because there is a boundary-layer diverter no incremental correction is required for the interaction of thepre-entry pressure gradient and the boundary layer and, as in Equation (5.10),

.

TABLE 9.1 Drag coefficient for scoop inlet

0 0.2 0.4 0.6 0.8 1.00 0.366 0.731 1.097 1.462 1.828

kf ksp sc ( CDf full)sh 0.308 0.193 0.111 0.051 0.012 0CDf full 0.300 0.300 0.300 0.300 0.300 0.300CDF 0.052 0.052 0.052 0.052 0.052 0.052CDd 0.034 0.034 0.034 0.034 0.034 0.034

CDsc 0.694 0.945 1.229 1.534 1.860 2.214

Ad wd dd=

0.184 0.014 0.0026= =CDd 0.25

0.00260.0190---------------- 0.034= =

m /m 02Ksc m /m 0( )

CD

PHtPH0----------

sc full

PHt TuPH0

---------------- d1u

PHt TlPH0

-------------- d1l

1d1u d1l( )

-------------------------=

/h /d1u = 0.1/0.176 0.568= = /h /d1l = 0.1/0.02 5.0= =

PHtPH0----------

sc full0.952 0.176 0.805 0.02{ } 10.176 0.02( )---------------------------------=

0.971=

P HtPH0------------ 0=30


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U. ESDU 86002The total pressure recovery ratio from Equation (5.11) is thus

,

.

Equation (5.12),

can be used to calculate the corresponding ram pressure efficiency and gives

.

Substitution of (PHt /PH0)sc into Equation (5.14) with ,

,

shows that the inlet is not choked for .

PHtPH0----------

sc

PHtPH0----------

sc full

PHtPH0------------+=

0.971 0+ 0.971= =

sc1 0.2M2+( )7/2 PHt /PH0( )sc 1

1 0.2M2+( )7/2 1---------------------------------------------------------------------------=

sc 0.916=

M 0.8=

m chm 0( )t

-------------

m chAt/A1( )m 0

-------------------------

0.579M

-------------

PHtPH0----------

sc

1 0.2M2+( )3 1.0086== =

m /m 0 At/A1( )1.0086 0.0181/0.0190( )1.0086 0.961==31


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U. ESDU 860029.1.4 Summary

The variation of the drag coefficient and ram pressure efficiency are plotted against mass flow ratio inSketch 9.1, which also illustrates the inlet dimensions derived during the calculations.

Sketch 9.1 Characteristics of scoop inlet

m

m0

0.0 0.2 0.4 0.6 0.8 1.0

sc

0.0

0.3

0.5

0.8

1.0

00

.

.

m

m0

0.0 0.2 0.4 0.6 0.8 1.0

CD sc

0.0

0.5

1.0

1.5

2.0

00

.

.

Operating point

0.600

0.060

0.1560.1760.190

0.020 0.006Dimensions in feet

0.152

0.01432This page Amendment D

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U. ESDU 860029.2 Flush Inlet Designed for Maximum Efficiency

9.2.1 Inlet size

For a curved-divergent ramp inlet operating at maximum efficiency, the size of the inlet necessary to capturea given mass flow is determined by calculating the mass flow parameter and using Figure 20to obtain .

For the inlet being considered

,

and for an aspect ratio w/dt = 4, Figure 20 gives

,

which implies

dt = 0.103 ft ,

and an inlet width w = 4dt = 0.412 ft.

As the lip profile is elliptical with ll = t = 0.25dt ,

ll = 0.0258 ftand t = 0.0258 ft.

The maximum external height of the inlet dmfl and the lip highlight height d1fl , measured in the inlet planerelative to the ramp floor, are then

The inlet capture area A1 is then d1fl w = 0.113 0.412 = 0.0466 ft2 .

ft,

and

ft.

m /0V02/dt

m

0V02------------------

0.01600.890 10 3 995 0.012--------------------------------------------------------------- 180.7= =

dt---- 0.097=

dmfl dt t ll tan+=0.103 0.0258 0.0258 7tan+=

0.126=

d1fl dt 0.5t ll tan+=0.103 0.5 0.0258 0.0258 7tan+=0.113=33


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U. ESDU 860029.2.2 Inlet drag

The overall drag coefficient is given by Equation (6.3),

.

(a) For the ram drag component, the first term on the right-hand side, the value of Kfl is obtained fromEquation (6.1),

.

For a ramp angle , a curved-divergent ramp and M = 0.8, Figure 10a gives

.

The momentum flow ratio is obtained from Figure 2 with andM = 0.8,

.

As a function of mass flow ratio the ram drag component is as follows.

(b) For the spillage drag coefficient, the second term, the factor allows for the ramp angle and isobtained from Figure 12. For ,

.

The factor kM allowing for the Mach number is obtained from Figure 13. For M = 0.8,

kM = 1.7 .

The variation of ksp fl with mass flow ratio is obtained from Figure 14. For M = 0.8 the variation is givenbelow.

0 0.2 0.4 0.6 0.8

0 0.315 0.630 0.944 1.259

0 0.2 0.4 0.6 0.8

1.0 0.54 0.21 0.03 0

CDfl 2Kflm

m 0------ kkMksp fl CD fl CD+ +=

Kfl kT0-------

=

7=

k 1=

T /0 /h /d1fl = 0.1/0.113 0.885= =

T0------- 0.787=

m /m 02Kfl m /m 0( )

k 7=

k 1=

m /m 0ksp fl34


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U. ESDU 86002The spillage drag coefficient at zero mass flow, , is obtained from Figure 11 as a function of(dmfl d1fl)/ll . For the elliptical lip

.

For a curved-divergent ramp

fl = 0.159 .

The variation of the spillage drag coefficient is then as follows.

(c) The incremental drag correction, the third and last term, is obtained from Figure 15 as a function ofmass flow ratio. For a curved-divergent ramp at M = 0.8, Figure 15b gives the following.

(d) The total drag coefficient as a function of mass flow ratio is obtained by summing the componentsestimated above, as set out in Table 9.2.

9.2.3 Pressure recovery

From Figure 17 the maximum ram pressure efficiency, , for an inlet with and w/dt = 4 is obtainedas a function of .

With , Figure 17 gives

.

The value of the modified mass flow ratio, , at this maximum efficiency is obtained from

0 0.2 0.4 0.6 0.8

0.270 0.146 0.057 0.008 0

0 0.2 0.4 0.6 0.8

0 0 0 0.010 0.110

TABLE 9.2 Drag coefficient for flush inlet

0 0.2 0.4 0.6 0.8

0 0.315 0.630 0.944 1.259

0.270 0.146 0.057 0.008 0

0 0 0 0.010 0.110

CD fl 0.270 0.461 0.687 0.962 1.369

CD

dmfl d1flll

-----------------------

0.5tll

--------- 0.5= =

CD

m /m 0k kM ksp fl CDfl

m /m 0CD

m /m 0

2Kfl m /m 0( )k kM ksp fl CD fl

CD

m 7=/dt/dt 0.097=

m 0.777=

m d1fl /m 0dt=35


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U. ESDU 86002Figure 18, with , as

,

or .

As a check, this corresponds to

the design value.

Figure 19 gives incremental changes that allow for the effect of parameters other than . The generalexpression for is given by Equation (6.5),

.

The correction for changes in the modified mass flow ratio, , is obtained from Figure 19a as a functionof

.

The following values are obtained for the present inlet.

Since the inlet has the optimum geometric configuration, and w/dt = 4, it follows from Figures 19band 19c that the corrections for ramp angle and aspect ratio are and .

Therefore the overall ram pressure efficiency is as set out in Table 9.3.

= 0.387 0.890 103 995 0.0466

= 0.0160 slug/s ,

0 0.2 0.4 0.6 0.8

0.425 0.205 0.013 0.234 0.453

0.101 0.040 0.001 0.037 0.050

/dt 0.097=

m 0.425=m

m 0------ 0.425

dtd1fl-------- 0.425 0.1030.113

------------- 0.387= = =

m 0.3870V0A1=

/dtflfl m mf w+ + +=

mf

md1fldt

--------

m

m 0------

0.387 0.113

0.103-------------

m

m 0------ 0.387 ==

m /m 0 m

mf

7= 0= w 0=36


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U. ESDU 86002

Substitution for in Equation (6.7),

,

with ,

shows that the inlet chokes at about , well above the design operating mass flowratio of 0.387.

9.2.4 Summary

The variation of the drag coefficient and ram pressure efficiency are plotted against mass flow ratio inSketch 9.2, which also illustrates the dimensions derived during the calculations.

TABLE 9.3 Ram pressure efficiency for flush inlet

0 0.2 0.4 0.6 0.8

0.777 0.777 0.777 0.777 0.777

0.101 0.040 0.001 0.037 0.050

0 0 0 0 0

0 0 0 0 0

0.676 0.737 0.776 0.740 0.727

m /m 0m

mfw

flfl

m chm 0( )t

-------------

m chAt/A1( )m 0

-------------------------

0.579M

------------- 1 0.2M2+( )3 fl1 fl( )

1 0.2M2+( )7/2-----------------------------------+

= =

At/A1 0.103 0.412( )/0.0466 0.911==m ch/m 0 0.86 0.87=37


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U. ESDU 86002

Sketch 9.2 Characteristics of flush inlet designed to operate at maximum ram pressure efficiency

m

m0

0.0 0.2 0.4 0.6 0.8 1.0

f l

0.0

0.3

0.5

0.8

1.0

00

.

.

m

m0

0.0 0.2 0.4 0.6 0.8 1.0

CD fl

0.0

0.5

1.0

1.5

2.0

00

.

.

Operating point

0.0258

0.1030.1130.126

7

Dimensions in feet

0.025838


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U. ESDU 860029.3 Flush Inlet Designed for Full Mass Flow

9.3.1 Inlet size

A first approximation to the size of the curved-divergent ramp inlet designed to capture 0.0160 slug/s whenoperating at full mass flow is obtained by setting

ft2 .

As the lip profile is elliptical

.

With and a throat aspect ratio of 4,

From Figure 1 with and ,

.

Thus using Equation (6.6)

If is increased to

ft2 ,

then ft.

For this value of , and Figure 1 gives .

,

and so = 0.0643 ft.

slug/s.

A1m

0V0------------

0.01600.890 10 3 995-------------------------------------------- 0.0181= = =

d1fl dt 0.5t ll tan+=

ll t 0.25dt= =

A1 w dt 0.5t ll tan+( )=4d2t 1 0.125 0.25 7tan+( )=

dt

/dt 0.1/0.0643 1.555= = M 0.8=

mT

m 0--------- 0.795=

m 0V0A1m Tm 0-------

=

0.890 10 3 995 0.0181 0.795=0.0127=

A1

A1 0.01810.01600.0127---------------- 0.0228= =

dt 0.0722=

dt /dt 1.385= m T / m0 0.810=39


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U. ESDU 86002For full mass flow Equation (6.6) now gives

This is a little too high, but three repetitions of the preceding numerical procedure quickly lead to the valuesA1 = 0.0224 ft

2 and dt = 0.0717 ft, for which Equation (6.6) predicts the desired full mass flow

slug/s.

At the design operating point the mass flow ratio is

.

For the elliptical lip

9.3.2 Inlet drag and pressure recovery

With the inlet dimensions determined, the calculation of CDfl and proceeds in an identical manner tothat set out in Sections 9.2.2 and 9.2.3, and is not given in detail. The results are summarised in Sketch 9.3,which also illustrates the inlet dimensions.

slug/s.

ft,

ft,

and

ft.

m 0.890 10 3 995 0.0228 0.810=0.0164=

m 0.0160=

m

m 0------

0.01600.890 10 3 995 0.0224------------------------------------------------------------------ 0.807= =

ll t= 0.25dt 0.0179= =

dmfl dt= t ll tan+0.0874=

d1fl dt= 0.5t ll tan+0.0785=

fl40


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U. ESDU 86002

Sketch 9.3 Characteristics of flush inlet designed to operate at full mass flow

0.0179

0.07170.07850.0874

7

Dimensions in feet

0.0179

m

m0

0.0 0.2 0.4 0.6 0.8 1.0

f l

0.0

0.3

0.5

0.8

1.0

00

.

.

m

m0

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00

.

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U. ESDU 860029.4 Comparison of Inlet Characteristics

To compare the drags and ram pressure efficiencies of th

esdu inlets characteristics

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