DERIVATION OF AERODYNAMIC THRUST CORRECTION FOR AN

25TH INTERNATIONAL CONGRESS OF THE AERONAUTICAL SCIENCES

1

AbstractThis “first principles” methodology enables thedirect “in situ” measurement and calculation ofthe wind tunnel drag effects acting upon theengine and it’s support structure, that arenecessary to derive the aerodynamic thrustcorrection for an indoor test facility, and thus,eliminate the need for outdoor “free field”testing.

1 Introduction and BackgroundAn indoor gas turbine engine test facility is bydesign, a low speed wind tunnel. The windtunnel effect is created by a secondary demandfor ambient airflow as a result of the ejectorpump action of the engine exhaust jet plumeentering the test facility exhaust collector(detuner). This ejector can create a secondarydemand for entrained/by-pass airflow that is upto five times greater than the initial airflowbeing demanded by the engine/intake. In thiscase, this would describe a test facility with anentrainment ratio of 5 to 1.

The creation of this wind tunnel effect isnecessary to assist expelling all undesirable hotgasses from the test cell, to enable meaningfulengine performance measurement andrepeatability in a stable and consistentaerodynamic environment of non-turbulentambient airflow. This will help eliminate anypotential instability, hot gas re-ingestion orvortex formation. Also, exposed elements of testfacility instrumentation/measurement systemscan be cooled with ambient airflow to avoidoverheating.

However, this wind tunnel effect createsa drag force acting upon the engine and itssupport structure, in an opposing direction to thetest facility thrust measuring load cells.Therefore, it is necessary to account for thisthrust drag debit (typically between 1-8%) withsome form of calibration, to enable measurednet thrust to be corrected to a set of referencedatum conditions that include still air (ISA sealevel static), to obtain a corrected gross thrust.

Traditionally, this calibration has beencarried out as a direct empirical back-to-backengine performance comparison between theindoor test facility and an outdoor free-field testfacility, which has an assumed infiniteatmosphere of still air. Unfortunately, due to aninconsistent UK climate and environmentalissues such as pollution and noise, testingengines on an outdoor test facility has becomelimited, time consuming and costly, withinconsistent results. Also, the recent generationof large civil engines has outgrown the currentlyavailable “industry standard” outdoor testfacilities, with evidence of ground effects andmicro-climates being confirmed usinginvestigative instrumentation and CFDmodelling. A measurement uncertainty ofapprox. ±0.5% (random) plus –0.5% to –1.0%(systematic) in gross thrust is currently beingestimated in these circumstances.

DERIVATION OF AERODYNAMIC THRUST CORRECTIONFOR AN INDOOR GAS TURBINE ENGINE TEST FACILITY

USING THE “FIRST PRINCIPLES” ANEMOMETER METHODNeil Parfitt Maurice BristowRolls-Royce plc Bristol Rolls-Royce plc Derby

Company Specialists within RR Performance Engineering

DERIVATION OF AERODYNAMIC THRUST CORRECTION FOR AN INDOOR GAS TURBINE ENGINETEST FACILITY USING THE “FIRST PRINCIPLES” ANEMOMETER METHOD

2

As an alternative to the free fieldmethod, the following first principlesmethodology has been developed and patentedto derive the aerodynamic thrust correction foran indoor test facility “in situ”, using arrays ofanemometers and static pressures in anextensive aerodynamic survey. This stand-alonemethodology can be used in isolation or withreference to any other source, effectivelyrelating an indoor test facility to free field usingfirst principles.

The following Fig 1 illustrates a typicalindoor “U” shaped sea level test facility withmain features, airflow paths and an introductionto first principle measurement planes andpositions.

Fig. 1. Typical Indoor “U” Shaped Sea LevelTest Facility

2 Basic MethodThe basic first principles method sums threedrag forces that are opposing thrustmeasurement, namely inlet momentum, cradleand base drags. Further detail regarding themeasurement apparatus, instrumentationrequirements, calculations and assumptionsassociated with this method are contained in thefollowing sub sections.

This basic method assumes that eitherthere is no engine debris guard being used, or ifthere is, that it is attached to the live and movingpart of the thrust measurement system.

In the event that a test facility has beendesigned with an engine debris guard that isfixed to the ceiling or floor and not attached tothe live and moving part of the thrust system,pressure loss through the debris guard willreduce the inlet momentum drag felt by theengine, and hence increase measured thrust. Inthis case, it will be necessary to derive a 4th termas a drag “debit” by removing the debris guardand carrying out empirical back-to-back testing.This is necessary because the first principlesmeasurement plane for inlet momentum drag ispositioned upstream of this pressure loss plane.It is a future aim to derive this term using apressure loss versus drag characteristic for afamily of engines on a test facility that use thesame debris guard.

2.1 Inlet Momentum DragThis is calculated from measuring the meanairflow approach velocity within theengine/intake “stream tube” using up to 9shrouded anemometers (1-9) positioned incruciform formation, axially positioned 2-3 air-meter throat diameters upstream of the frontface of the intake bell mouth flare (see Fig 2).

This axial position is chosen for 2reasons :-• To only capture the test facility wind tunnel

effect before any acceleration into the intakebell mouth flare occurs.

• To ensure that there is sufficient distance toallow any wakes from the measurementapparatus to re-mix, and hence not affectengine stability or performance.

The radial positions specified on Fig 2ensure that at least 1.5 × air-meter throatdiameters of approaching stream tube iscaptured for measurement.

EJECTOR INDUCED ACCELERATING ENTRAINED AIRFLOW

ENTRAINED/BY-PASS AIRFLOW

FLARED BELLMOUTH VENTURI AIRMETER AND DEBRIS GUARD

ENGINE HOT CONVERGING EXHAUST NOZZLE

PYLON

MOVING PART OF THRUST SYSTEM

FIXED PART OF THRUST SYSTEM

THRUST MEASURINGLOAD CELL

TEST FACILITY INTAKE SYSTEM

TEST FACILITY EXHAUST SYSTEM

o o o o o o o oo o o o o o o oo o o o o o o oo o o o o o o oo o o o o o o oo o o o o o o oo o o o o o o oo o o o o o o o

DETUNER

AIRMETER STREAM TUBE AIRFLOWX

X

X

X

X

FORWARD ANEMOMETER PLANE

X XXXX

NOZZLE STATIC PRESSURE POSITIONS

X XX

X

CRADLE ANEMOMETER POSITIONS

EJECTOR INDUCED ACCELERATING ENTRAINED AIRFLOWEJECTOR INDUCED ACCELERATING ENTRAINED AIRFLOW

ENTRAINED/BY-PASS AIRFLOWENTRAINED/BY-PASS AIRFLOW



PYLON







DETUNER





PYLON




PYLON









DETUNER


DETUNER

AIRMETER STREAM TUBE AIRFLOWAIRMETER STREAM TUBE AIRFLOWX

X

X

X

X


X

X

X

X

X


X XXXX


X XXXX


X XX

X


X XX

X



3

Fig. 2. Upstream Anemometer Positions(Cruciform Formation)

Note:• The 4 off additional anemometers (10-13) shown in

Fig 2 are used for an overall aerodynamic stabilitysurvey and not these specific inlet momentum dragcalculations.

In the case of engines with an air-meterthroat diameter less than 1,5 meters, 5anemometers are likely to suffice with the 0.375throat diameter radius position anemometers (2-5) not being required.

This method enables the calculation of abasic area weighted average of airflow velocitythat is considered acceptable for manyapplications. Any variation in airflow approachvelocity profile about the 5 or 9 anemometersgreater than ±15% from mean, is consideredunacceptable.

Therefore Intake Momentum Drag (1)= (W1 × V0)/1000 kN

Where:-W1 =Observed engine inlet mass airflow - kg/sV0 =Mean velocity of 1-5or9 anemometers - m/s

Note:• Engine inlet mass airflow should be measured in a

calibrated air-meter to within an accuracy of ±1.0%.

Fig. 3 illustrates a typical example oftubular cruciform apparatus to locate theseforward anemometers.

Fig. 3. RB199 Engine at RAF Marham, UK

Alternatively, a similar but less intrusivemethod to locate the forward anemometersusing pairs of taught wires can be seen on Fig 4.

Fig.4. Tay Engine at RR Dahlewitz,Berlin, Germany.

C

2

D A

B

Ai r - meter Anemometer

Test Cell Boundary

Axes through engine centre - line

Round section tubing of no more than 2” diameter

10

13

12

1

8

6

4

7 3 11 5 9

Note. Anemometers to be Symmetrically positioned A bout Cruciform as follows : - A = 0.375 × A ir - Meter T hr oat Dias (1 - 5) B = 0.750 × A ir - Meter T hr oat Dias (1, 6 - 9) C = D (6 - 13)


4

As a further alternative to the cruciform type ofapparatus, a vertically traversing horizontalboom is currently being used to locate theforward anemometers in some of the large civilengine test bed applications.

This enables a much greater array of airflowvelocity measurement in a larger arena, leadingto a better understanding of inlet stream tubeprofile shape and the ability to carry outpreferred mass momentum flow weightedaveraging, using contour plotting andspecialised software.

Figure 5 highlights a typical traversing boommeasurement array with the basic cruciformformation and an approximated stream tubesized for a Trent 900 engine, shown forcomparison.

Fig. 5. Typical Upstream AnemometerPositions (Traversing Boom)

Figure 6 illustrates a typical traversing boominstallation to locate the forward anemometers.

Fig. 6. Trent 900 Engine at RR Derby

One further advantage of using this traversingmethod is that a much better understanding of“full” test facility airflow profiles and hencefuture design capability can result from thisgreater array of measurement. This can also beused to support CFD modelling.

Typically, this inlet momentum termwill account for between 70%-90% of the totalthrust correction and hence, is by far the largestterm.

If it is necessary to include the “fixed”debris guard drag debit term, this can typicallyreduce this inlet momentum drag term by up to50%.

2.2 Cradle DragThis is derived by calculating the pressureloading of entrained/by-pass cell airflow actingupon the frontal blockage areas of all live andmoving parts of the test facility thrust cradle andattached obstructions. This pressure loading iscalculated by measuring the airflow velocitiesadjacent to (approx. 10 cm from anemometercenter line to blockage component edge) thesefrontal blockage areas, using up to 10 shroudedanemometers evenly spread around the thrustcradle and attached obstructions.

13 nominal Traversing Boom Height Positions in metres

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CL S5 S4 S3 S2 S1 S6 P6 P5 P4 P3 P2 P1

1

2

4

3

10

9

8

7

6

5

11

12

13 Test Cell Boundary

View looking from front of Test Bed towards Detuner

Airmeter Bellmouth Flare 13 off Anemometers at 1 metre spacing

Axes through Engine Centre-Line

Vertical Traversing Boom

Side Rails


5

Deciding where to position theseanemometers involves an element of judgment,since no two test facilities are likely to beexactly the same. If it is possible to achieve areasonable symmetrical average spread, thencalculated mean averages can be used asdepicted in the following equation (2).However, if this is not possible or significantlydiffering cell airflow velocities are being seenaround the thrust cradle, then individualanemometer values in conjunction with theirrespective adjacent blockage areas can besummed after being broken into separatecomponents.

Therefore Pressure Loading (∆p) (2)= P0 1 - ( 1 ) kPa

(1 + (6.0449 × 10-6 × VCELL2))

Where:-P0 = Cell static pressure – kPaVCELL=Mean vel of up to 10 anemometers – m/s

Notes:• The preferred position for cell static pressure

measurement is approx 1 meter outboard from thetest cell wall in the axial plane of the cold/hot nozzlesat engine centre line height.

• The constant contained in equation (2) above fixesthe cell static temperature at an ISA day value inconjunction with the speed of sound. This is done toeliminate the need to measure and calculate local cellstatic temperatures that would be required for thetraditional Bernoulli equation (½ρV2). Themeasurement uncertainty associated with this hasbeen approximated as ±0.01% in gross thrust for anambient temperature change of ±20 K on EJ200 andis considered negligible.

The total frontal geometric blockagearea of all moving elements of the test facilitythrust cradle and attached obstructions must bemeasured. If any blockage elements are shieldedby further upstream blockage, then it isrecommended that they should not be countedunless they are greater than 5 obstruction widthsdownstream of the upstream blockage. Also, itis recommended that flexible components suchas hoses, cables and tubes should not becounted, because any potential drag effects arelikely to be negligible.

A mean Cd that relates to this total thrustcradle blockage area needs to be calculated. Therespective shape Cd’s can be acquired byreference to Hoerner’s Fluid Dynamic Drag –Chapter V111 (interference drag).

Fig 7 illustrates some typical referenceCd values used :-

Fig. 7. Example Reference Cd’s to be used forVarying Thrust Cradle Blockage Area Shapes

Therefore Cradle Drag (3)= ∆p × ACRAD × CdCRAD kN

Where:-∆p = Pressure loading as equation 2 - kPaACRAD = Meas total frontal blockage area – m2

CdCRAD = Calculated mean blockage area Cd

1.16

1.17

SHAPE Cd

1.55

1.55

2.00

2.20

1.60

1.98

Airflow

1.20

2.30

2.05


6

Two examples of anemometer positions,measured cradle blockage area and calculatedrespective Cd’s are shown on Figs 8 & 9respectively

Fig. 8. EJ200 Engine Floor MountedThrust System at RR Bristol

Note:• Geometric Square Cd

Blockage MetresAreaA-ETotal =(0,8628) Mean =(2.00)F-K Total =(1,5272) Mean =(1.87)L-OTotal =(0,645) Mean =(1.92)Overall Total =(3,0350) Overall Mean =(1.92)

Fig. 9. Trent 800 Engine Overhead MountedThrust System at RR Derby

Note:• Geometric Square Cd

Blockage InchesAreaA-ETotal =(6942) Mean =(1.81)F-K Total =(17832) Mean =(2.02)L Total =(1056) Mean =(1.60)Overall Total =(25830) Overall Mean =(1.95)

Typically, this cradle drag term is likelyto account for between 5-25% of the total thrustcorrection.

Whilst this term is associated with testfacility hardware only, if it is judged that someengine or slave test hardware componentsconstitute potential additional drag, then theyshould be counted as a separate drag usingsimilar techniques and calculations as used forthis drag term.

N

O

A

11 A

B

F H

14 13 H

12

C

G

O


Anemometers

B Engine Centre Line (viewed from front)

D

16

17

I

I K

K

M

J

L

L 15

E

10

E

K

G F

F

ZTCRDL4

Anemometers

Fixed Beam

H

H

H ZTCRDL6

ZTCRDL3

ZTCRDL1

H

I

I I


ZTCRDL5

ZTCRDL2 Anemometers

K

K

J J

ZTCRDU2 A D

D D ZTCRDU3 B

C

D ZTCRDU1

E

Anemometers

Engine centre line (viewed from front)

L


7

2.3 Base Drag (Detuner Suction)This is derived from calculating the mean staticpressure depression as a result of acceleratingentrained/by-pass airflow velocities as they flowover the convergent exhaust nozzle to enter thedetuner. The ejector creates a suction force thatcan be calculated using ∆p × area change. This∆p is calculated from measuring the mean staticpressure from at least three equal axial planes(front, mid, rear) along the outer skin of theconical section of the exhaust nozzle, with thefinal position being measured by at least 4circumferential sensors to evaluateprofile/ejector effectiveness in an overallaerodynamic survey (see Fig 10).

Fig. 10. Typical Convergent Nozzle StaticPressure Arrangement Configuration

Notes:-• Static pressures to be positioned as shown with the

end turned to be normal to the airflow by no morethan approx. 7mm.

• Mean converging nozzle static pressure equates to themean of the 3 planes F,M&R, noting that the stationR will be the mean of at least 4 sensors R1-R4.

A much greater array of static pressuremeasurement can be considered for more criticalor complex shape applications.

Therefore Base Drag (4)= (P0 – P7.5) × (A7 – A8) kN

Where:-P0 = Cell static pressure (as equation 2) – kPaP7.5 = Mean conv nozzle static pressure – kPaA7 = External front conv nozzle area – m2

A8 = External rear conv nozzle area – m2

Fig 10 illustrates a typical single nozzleconfiguration for a turbojet or mixed turbofan.In the case of a non-mixed separate jetturbofans, where detuner entrained cell airflowis not in direct contact with the hot exhaustnozzle skin, this term only applies at theforward cold nozzle position.

If the test facility detuner is a telescopicdesign, then it is possible to optimise the ejectoreffect and hence, reduce local airflow velocitiesand thrust drag terms to a level that isaerodynamically preferable overall. Thisincludes ensuring that no hot gas is spilled intothe test cell that can be re-ingested by theengine, which can be measured during anoverall aerodynamic survey.

Two examples of nozzle static pressureapplications are shown on Figs 11 & 12respectively.

Fig. 11. EJ200 Mixed Single Nozzle StaticPressure Application at RR Bristol

6 off Static Pressure Sensors, made from basic 1,6mm O/D stainless/steel tubing.

External Front Converging Nozzle Area

External Rear Converging Nozzle Area

R1 M1

F1

R3

R2

R1

R4 Converging Engine Exhaust Nozzle

View on “A”

“A”


8

Fig. 12. Trent 900 Separate Jet Cold NozzleStatic Pressure Application at RR Derby

2.4 Total Drag (Overall Aerodynamic ThrustCorrection)This is simply the sum of the former threecomponents and can be converted into anoverall aerodynamic thrust correctioncharacteristic, which can be deployed in aperformance pass-off or analysis softwareprogramme, preferably as a correlation againstmeasured thrust or inlet flow function.

Typically, this total drag term is likely tobe in the region of 1-8% of gross thrust,depending on engine size in relation to it’s testfacility.

A measurement uncertainty of less than±0.25% in gross thrust is currently estimated forthis method. This will increase to approx.±0.30% if the test facility has a “fixed” debrisguard that necessitates deriving the additional4th component previously discussed (debrisguard debit drag).

2.5 Basic Method AssumptionsThe basic method assumes a static pressure fieldequilibrium throughout the test cell and hence“control volume” thrust momentum box (whichincludes the forward anemometer plane) – seeFig 13.

This assumption is consideredacceptable for test cell applications in which theairflow velocity is stable, uniform and in theregion of 30 ft/sec or less, with an entrainmentratio greater than 2 to 1 and an overallaerodynamic thrust correction of 4% or less. Inthese cases it is believed that any additional orunaccounted forces, particularly P0(A0-A1) andP0(A1-A8) shown on Fig 13 are considerednegligible (<0.1% of gross thrust), and thereforecancel in the following equation (5).

Fig. 13. Current Basic “First Principles” MethodControl Volume Assumptions

Notes:• The potential force P0.(A1-A8) represents any

buoyancy forces acting on the engine due to thereduction in area (front to rear), assuming all externalstatic pressure is P0

• The potential force P0.(A0-A1) has two parts:1) The force acting upon the rear of the bellmouth

flare2) The axial momentum lost by the engine stream

tube flow, due to the turning of the stream lines

Therefore Gross Thrust (5)= W8.V8+P8.A8 − P0.(A1-A8) − P0.(A0-A1 ) +(W1.V0+P0.A0) + Cradle Drag + Base Drag

= W8.V8 − (P8-P0)A8 + W1.V0 + Cradle Drag+ Base Drag

= Net Thrust (measured) + Inlet Momentum +Cradle Drag + Base Drag

INLET MOMENTUM

BASE DRAG

W1.V0+P 0.A0

PYLON

A0 A1

P 0.(A0 - A1)

W8.V8+P8 .A8

P0.(A1 - A8 )

CRADLE DRAG

A8 NET THRUST


9

3 Revised Full MethodCurrent industry standard large civil engineindoor test facilities are unlikely to be able toachieve the aerodynamic requirementassumptions listed for the basic method,particularly regarding minimum airflowvelocity. This, coupled with the likelihood of atighter uncertainty assessment requirement forcustomer compliance demonstration,necessitates a far more rigorous assessment ofany potential additional and unaccounted forcesthat are deemed negligible with the basicmethod.

Therefore, the additional terms deemednegligible and shown on Figure 13 are likely tobecome more significant and need to bequantified as additional engine buoyancy forces.Figures 14A & B illustrate a revised full controlvolume thrust momentum box in descriptionand equation form respectively, that enable firstprinciples thrust correction to be aligned withother methods (“ideal free field” or “nozzlerig”), and thus “thrust in flight” accountingduring civil engine compliance demonstration tothe customer.

Fig. 14A. Revised Full “First Principles”Method Control Volume Assumptions

Note:• Pre-entry streamtube force (including rear bellmouth

scroll buoyancy force) can be estimated fromentrainment flow, = Entrainment momentum (1) +Entrainment buoyancy force - Entrainmentmomentum (0)

Therefore Gross thrust (6)= Net thrust + Inlet momentum - Pre-entrystreamtube force – External b/m scrollbuoyancy + Engine buoyancy forces + Cradledrag + Friction drag - Nozzle exit buoyancyforce

Fig. 14B. Revised Full “First Principles”Method Control Volume Assumptions

Note:• ∫ (p-p0) δAstream (including rear bellmouth scroll

buoyancy force) = We1.Ve1 + Ae1(pe1-pe0) -We0.Ve0

Therefore Gross thrust (7) = FN + W0.V0 - ∫ (p-p0) δAstream - ∫ (p-p0)δAbm + ∫ (p-p0) δAeng + Cradle drag +Friction drag - Aj(pref-p0)Where:-pref = pj * (p0 “free-field” / pj “free-field”) for“free-field” nozzle coefficient accounting, andfrom entrainment flow

The following additional requirements (relativeto the basic method) need to be considered inorder to enable the derivation of the additionalterms :-• Derive the stream tube pre-entry force using

CFD model integration or simple 1Dentrainment flow calculation (not directlymeasurable).

• Derive the bell mouth buoyancy (pull-off)term also using the CFD or 1D calculation asabove.

• Fit static pressure measurement sensors (p0)at the traversing boom (A0) plane.

• Fit static pressure measurement sensors tothe rear face of the bell mouth scroll.

• Fit static pressure measurement sensors (pj)at the (Aj) plane. This can be achieved usingan existing design of piezo-ring, which isadditional to current base drag static pressuremeasurement.

Engine

Cradle drag

��

Inlet momentum

Pre-entry streamtube force

DetunerFriction drag Engine buoyancy

forces (base drag)

Net thrust + nozzle exit buoyancy force

��

��

Boundary layer

Entrainment flow momentum + buoyancy force Entrainment flow momentum +

buoyancy force

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External b/m scroll buoyancy

Engine

Cradle drag

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Inlet momentum

Pre-entry streamtube force

DetunerFriction drag Engine buoyancy

forces (base drag)

Net thrust + nozzle exit buoyancy force

��

��

Boundary layer

Entrainment flow momentum + buoyancy force Entrainment flow momentum +

buoyancy force

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External b/m scroll buoyancy

Engine

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A0p0

A1

Aj

W0.V0

∫ (p-p0) δAstream

DetunerFriction drag ∫ (p-p0) δAeng

FN + Aj(pref-p0)

pj

��Boundary layer

We0.Ve0 + Ae0(pe0-p0) We1.Ve1 + Ae1(pe1-p0)Ae1pe1

Cradle drag

P2

Ram ratio is defined As P2 / pref

∫ (p-p0) δAbm

Engine

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��

A0p0

A1

Aj

W0.V0

∫ (p-p0) δAstream

DetunerFriction drag ∫ (p-p0) δAeng

FN + Aj(pref-p0)

pj

��Boundary layer

We0.Ve0 + Ae0(pe0-p0) We1.Ve1 + Ae1(pe1-p0)Ae1pe1

Cradle drag

P2

Ram ratio is defined As P2 / pref

∫ (p-p0) δAbm


10

From this information, the following additionalterms can now be quantified :-• Pre-entry stream tube force (turning stream

lines) from CFD and/or 1D calculation.• Pre-entry stream tube force (bell mouth pull-

off) from CFD and/or 1D calculation and/ormeasured static pressure ∆p (bellmouth –boom (p0)).

• Nozzle exit buoyancy force from CFD and/ormeasured static pressure ∆p ((pref) – boom(p0)).

• Engine buoyancy forces shown are currentlyrepresented by base drag in the basic method.

• Friction drag is considered negligible.Notes:-• pref is determined from pj with a correction defined

from free field or noxzzle rig.• pref is also used for ram-ratio definition corrections

in the engine performance synthesis model.• p0 should now replace cell pressure as the datum for

basic method cradle and base drags.

Some of the above terms are self-canceling.However, the net result provisionally defined onRR’s highest flowing engine in its master testfacility (Trent 900) was between –0.2% & -0.3% of gross thrust relative to the basic firstprinciples method.

It should be re-iterated that when testing smallerlow flow engines in the regime described for thebasic method, this is more likely to be in theregion of -0.02%.

4 Advantages and Benefits of using FirstPrinciples

• Significant Quality Improvements -Derived thrust correction uncertainty(including repeatability) improved from anestimated ±0.8% to ±0.25%. This alone hasled to improved customer satisfaction.

• Major Cost Benefits to EliminatingOutdoor Testing – Demonstrated duringsuccessful compliance demonstration for theGulfstream G450 aircraft programme (seealso reasons next bullet).

• Major Lead Time Benefits to EliminatingOutdoor Testing – Estimated to be in theregion of 3-6 months per installation. Thisincludes eliminating two of the calibrationlegs in a B-A-B type cross-calibration andthe time taken waiting for acceptableenvironmental conditions in a typicalinconsistent UK climate.

• Ability to Calibrate an Indoor TestFacility from “Day One” of a New EngineProject – This has previously taken up totwo years before the formal customercompliance demonstration is carried out,thus carrying significant risk well into theprogramme.

• Major Cost and Lead Time Benefits toEliminating Regular Indoor Test FacilityCross-Calibration – Ability to re-calibrateindoor test facilities (including masters) atany time “in situ”, without reference to anyother source, thus eliminating expensive A-B-A type back-to-back testing.

• Ability to Identify and Quantify TestFacility Aerodynamic Design Problems –The vast array of acquired measuredaerodynamic data is helping to identifyproblems from the outset and diagnoseinstant corrective action, thus avoiding thetraditional and more expensive timeconsuming culture of “try it and see” fixes.

• Ability to Support the Next Generation ofImprovements – The vast array of acquiredmeasured aerodynamic data is also enablingbetter CDF modeling of test facility design,with the potential future ability to be able topredict thrust correction accurately.

Documents

DERIVATION OF AERODYNAMIC THRUST CORRECTION FOR AN