Validation of turbine noise prediction tools with acoustic - MTU Aero ...

Validation of turbine noise prediction tools 

with acoustic rig measurements 

Dominik Broszat ∗ , Detlef Korte 

MTU Aero Engines GmbH, Department Aerodynamics, Aeroelasticity, Aeroacoustics Turbine 

Dachauer Strasse 665, D-80995 Munich, Germany 

Ulf Tapken 

German Aerospace Center (DLR), Institute of Propulsion Technology, Berlin, Germany 

Mathias Moser 

Graz University of Technology, Institute of Thermal Turbomachinery and Machine Dynamics, 

Graz, Austria 

Well-defined measurement data are of key importance for the validation and calibration 

of numerical methods. Accordingly, an acoustic test rig resembling the last stage of a low 

pressure turbine (LPT) of a modern turbofan engine has been established at the Graz 

University of Technology in Graz, Austria within the frame of the European Research 

Project VITAL. It has been designed to allow for a clear separation of the contributors to 

turbine noise by a systematic selection of the individual blade numbers of the subsequent 

rows. This setup permits the classification of the measured or computed tones with respect 

to their origin and source mechanism. Within the VITAL project, several acoustic test 

campaigns have been carried out: a datum (or reference) configuration as well as variations 

of the axial gapping between the stator, rotor, and the EGV. These give insight into the 

relative weighting of the above mentioned source mechanisms and the influence of distance 

variations on noise generation. 

Abbreviations: 

ADP Aerodynamic Design Point LPT Low Pressure Turbine 

BPF Blade Passage Frequency OP Operating Point 

EGV Exit Guide Vane PWL Sound Power Level 

EO Engine Order RMA Radial Mode Analysis 

FFT Fast Fourier Transform SNR Signal-to-Noise Ratio 

GTF Geared Turbofan STTF Subsonic Test Turbine Facility 

IGV Inlet Guide Vane TEC Turbine Exit Case 

LEE Linearized Euler Equations VITAL EnVIronmenTALly Friendly Aero Engine 

I. Introduction/ motivation 

Within the last decades, the noise radiated from an aircraft engine has been reduced drastically. This 

can be attributed both to source noise reduction and passive measures to reduce the noise radiated to the 

far field. The latter refers to a maximization of the acoustically lined area within the engine inlet and bypass 

duct. In parallel, both jet and fan noise have been reduced in the course of constantly increasing bypass 

ratios motivated by the maximization of the propulsion efficiency of the engine. 

∗ Corresponding author, PhD, email: dominik.broszat@mtu.de 

1 of 13 

American Institute of Aeronautics and Astronautics

However, due to these noise reductions of the main contributors to engine noise, other components have 

recently gained importance, e. g. the turbine (and especially the low pressure turbine (LPT)) within the 

rear arc. This is especially true in operating conditions where the noise floor of the other components is 

fairly low, e. g. at approach power (compare figure 1), such that the LPT can become the dominating noise 

source within a certain frequency range. Related to current engine developments, this fact gains even more 

importance in the case of geared turbofan (GTF) designs where fan noise can be reduced even more compared 

to conventional applications and thus the relative weighting of the LPT increases. 

(a) (b) 

Figure 1. Contributors to aircraft noise at takeoff and approach (schematic) [Ref. 1] 

Motivated by this growing importance of LPT noise in modern aircraft engines and for the validation of 

analytical and numerical tools used for the design and assessment of turbomachinery, an acoustic test rig 

has been established and put into operation at the Graz University of Technology in Graz, Austria within 

the European research project VITAL (EnVIronmenTALly Friendly Aero Engine). It has been carefully 

designed to allow for a clear separation of the individual noise contributors within a 1 1 

2 

stage LPT at 

minimum disturbances. In addition to the ability to determine the noise generation origins and the relative 

weighting of the individual interaction mechanisms in relation to the overall LPT noise for a reference 

configuration, several acoustic measurement campaigns have been conducted involving geometric variations 

as well as acoustic measures to reduce both the noise generation and its radiation from the LPT exit. 

Within this paper, the setup of the acoustic test facility, a short description of the measurement campaigns, 

the general features of the tools used for predicting the LPT noise, and a comparison of the experimental 

and numerical results will be given. 

II.A. General setup 

II. Design of the acoustic test rig 

In this section, only a brief description of the test rig setup will be given, a more detailed version is available 

in Ref. 2. Figure 2 shows a sectional drawing of the STTF (Subsonic Test Turbine Facility) located within 

the laboratory of the Graz University of Technology. As indicated by the arrow, the flow enters the rig 

through a spiral inlet casing in which it is turned into the axial direction of the rig. To cancel the remaining 

swirl, a deswirler unit has been included ahead of the turbomachinery section. 

The center part of the rig contains the 1 1 

2 

stage LPT, i. e. one turbine stage consisting of a stator and a 

rotor plus an inlet (IGV) and exit guide vane (EGV or TEC). For the ease of assembly and disassembly, the 

rotor is attached to an overhung-type shaft connected to a water brake to adjust the rpm of the LPT. By this 

design, both the turbomachinery parts and the downstream acoustic test section can be easily separated and 

modified. The corresponding airfoil counts of the four rows are listed in table 1. As mentioned above, they 

have been carefully chosen to allow for the clear determination of the noise generation origin of propagating 

acoustic modes. 

For the acoustic analyses, an array of flush-mounted microphones has been included on both the inner 

and outer annular duct walls downstream of the TEC. To reduce the number of sensors, three microphone 

plates with linear arrays have been positioned equally spaced on the circumference of a 360 ◦ rotatable duct 



Figure 2. General arrangement of STTF acoustic rig 

Table 1. Airfoil counts of STTF rig 

IGV S1 B1 EGV 

83 96 72 15 

section. A sectional drawing of this part of the rig and a closeup of one of the microphone plates can be seen 

in more detail in figure 3. 

(a) acoustic test section of STTF rig (b) microphone plate 

Figure 3. Sectional drawing of acoustic test section and closeup of microphone plate 



As can be seen, the microphones have been equally spaced in the axial direction and have been arranged 

in two lines to reduce the axial extent. Based on the assumption that the mean aerodynamic and aeroacoustic 

field are time-invariant for the duration of one test campaign, this setup using a traversable duct section 

permits the recording of the full circumference with a drastically reduced number of sensors. Finally, at the 

downstream end of the acoustic test section, the flow is redirected again into the vertical direction and leaves 

the laboratory through an exhaust stack at ambient conditions. 

Concerning the aerodynamic design, the LPT stage has been designed to represent the last stage of an 

LPT of a modern commercial aero engine at approximately half scale. With respect to the characteristic 

small expansion of the flow path at the rear stages of a typical engine, the cylindrical annular duct cross 

section motivated by acoustic reasons can be justified. Although designed for an aero design point (ADP), the 

rig has been operated at the three noise certification operating points (OP) approach, cutback, and sideline 

which are related to the ADP by the following factors (compare Table 2) to represent realistic operating 

conditions. 

Table 2. Rig operating points in relation to ADP [Ref. 2] 

parameter ADP Sideline Cutback Approach 

total pressure ratio [%] 100 97 85 73 

mass flow [%] 100 95 75 50 

shaft speed [%] 100 97 91 71 

shaft power [%] 100 87 48 16 

II.B. Aerodynamic and acoustic measurements 

The aerodynamic conditions within the test rig have been determined at several axial positions as indicated 

in the sketch in figure 4. In the inlet plane (plane 0) to the turbomachinery stage, the inflow is assumed to be 

circumferentially and radially uniform, and the total pressure has been recorded by a Kiel probe at 5 radial 

locations. Together with the total temperature, this value has been used to normalize the measurements in 

the downstream planes to cancel slight variations in the operating conditions (mainly due to variations of 

the inflow characteristics). Behind the IGV, in plane A, five-hole probe measurements have been performed 

covering 4 sectors of 24 ◦ azimuthal extension on the circumference. Within these sectors, 21 radial and 17 

axial positions have been recorded starting approximately 4 mm from the duct walls (due to the dimensions 

of the probe). 

Figure 4. Aerodynamic measurement planes in STTF rig 



Figure 5 shows an exemplary plot of 

the Mach number distribution in these 

sectors at approach. Apart from variations 

in the radial direction, the wakes of 

the IGVs can be clearly observed. 

In the downstream planes B and C 

(downstream of the rotor and TEC, respectively), 

only 1 pitch of the corresponding 

upstream stator row has been traversed. 

Results of these flow measurements 

can be found in Ref. 3. As will be 

mentioned later on, these aerodynamic 

measurements have been of great importance 

for the comparison with the numerical 

simulations. In addition, static pressure 

taps have been placed on the EGV 

and the inner and outer duct walls within 

the acoustic measurement section. 

Concerning the acoustic measurement 

and analysis process, the DLR (German 

Aerospace Center) has established a Figure 5. Mach number distribution in plane A, OP approach 

mode analysis technique which has been 

successfully applied to numerous rigs in the past years. 4, 5 It is based on the measurement of the instantaneous 

pressure at various axial and circumferential positions at the inner and outer duct walls and a modal 

decomposition into its circumferential and radial components. In order to achieve mode analysis results of 

high accuracy with a minimum measurement effort, the sensor configuration was optimized for the given test 

rig according to the guidelines given by Tapken and Enghardt. 6 At this point it should also be mentioned 

that all analysis effort within the framework of the VITAL project was focussed on tonal noise generated by 

the interaction of relatively moving blade rows at harmonics of the Blade Passage Frequency (BPF). 

During the measurements, the time series of the 72 traversable microphones have been directly recorded 

in a throughput mode on a hard disc at a sampling rate of 51.2 kHz (i. e. a usable frequency range up to ≈ 20 

kHz). With a circumferential step width of 2◦ this yields 4320 measurement points overall. The subsequent 

offline analysis process can then be split into 3 steps: 

• an adaptive resampling of the time series, 

• an FFT of these modified time series, 

• the Radial Mode Analysis (RMA) at the BPF harmonics. 

The first process involves a resampling of the time series using a rotor trigger signal to account for 

variations in the shaft speed. It impressively enhances the quality of the rotor-coherent signals by increasing 

the signal-to-noise ratio (SNR) of the relevant tones. The FFT (Fast Fourier Transform) then transforms 

the signals to the frequency domain in which the final step, the RMA can then be performed. The RMA 

decomposes the measured sound field at one selected frequency (here the BPF harmonics) into its azimuthal 

and radial modal constituents by a matrix inversion of the transformation matrix between the vector of the 

measured sound pressure levels (at all microphone positions) and the modal amplitudes of all propagating 

mode orders (m,n). These results can then be attributed to the individual acoustic interactions of the LPT 

blade rows and compared to the numerical results. However, at this stage, a decomposition of only the 

azimuthal mode orders (m) will be presented. 

III. Aeroacoustic prediction tools 

In this section, the two methods used for predictions (both pre- and post-test) will be briefly presented. 

However, for a more detailed description, previous publications will be referenced. The two tools are a semiempirical/ 

analytical predesign tool based on meanline aerodynamic data called TCLOW and a numerical 

3d CAA code based on the Linearized Euler Equations (LEE) and 3d aerodynamic flow fields called Lin3d. 



III.A. Semi-empirical/ analytical prediction tool 

This fast predesign tool takes into account the viscous wake interaction of relatively moving blade rows. 

Thereby these rows do not necessarily have to be directly neighboring but interactions with further upstream 

generated wakes are computed as well. The correlation used for the computation of the wake width and depth 

has been established by Kazin et al. for turbine applications and is also well-known as the GE correlation. 7 

Following, the unsteady aerodynamic forces, i. e. the blade loading, is calculated using a model by Kemp and 

Sears. 8 Finally, the acoustic power level (PWL) is computed using correlation formulas derived by Lowson. 9 

Subsequent steps in the TCLOW tool include the propagation through downstream stages (if any) using 

transmission and attenuation coefficients, scattering effects at downstream blade rows, and the radiation to 

far field microphone positions on a 150 ft. arc, the standard setup for static engine measurements. Concerning 

the propagation and radiation characteristics, TCLOW also includes an analytical haystack model to account 

for the broadening of tones due to the propagation through the engine shear layers. In addition, the PWL 

results are adjusted to a database of measured engines for an absolute calibration of the tool. 

III.B. Linearized Euler analysis 

The working principle of the time-linearized 3d Euler code Lin3d is explained in more details in a number of 

papers published within the recent years, e. g. Refs. 10, 11, and 12. However, at this point, it should be briefly 

mentioned that it computes the small acoustic harmonic perturbations using a linearized version of the 3d 

Euler equations around a steady state aerodynamic solution. Currently, this steady flow field is computed 

using a 3d RANS code especially adapted to turbomachinery applications (TRACE code), a modification to 

the previous procedure of using an Euler code for the CFD computation. This allows for the more realistic 

representation of the real flow phenomena within a turbomachine, but also introduces numerical problems 

due to the boundary layers and the zero velocity at the blade surface which cannot be processed within the 

Euler equations. 

Concerning the perturbations superimposed on the steady flow, both a viscous wake based on the correlation 

mentioned in the previous paragraph, 7 and a Fourier decomposition of the upstream or downstream 

static pressure field (on the upstream or downstream blade row) can be prescribed at the inlet or outlet of 

any cascade. Due to the relative motion of the neighboring blade rows, they have to be computed separately 

in their respective relative coordinate systems and are coupled at their boundaries. At these locations, 4 

characteristic variables, i. e. up- and downstream propagating pressure waves, and downstream propagating 

vorticity and entropy waves can be exchanged and non-reflecting boundary conditions are imposed. However, 

this is currently performed at various radial positions neglecting a true radial mode decomposition. The final 

result of these coupled computations are the modal PWLs at the boundaries of the computational domain 

which can be compared e. g. to acoustic measurements. 

It can be imagined that the convergence of this coupled unsteady problem can be hard to achieve, 

especially in cases where modes experience a cut-on/ cut-off transition along the annulus. In addition, 

the boundary layers of the airfoils and regions of separated flow in off-design points (as e. g. the approach 

condition) complicate the finding of a converged numerical solution. In these cases, several methods are 

available to counteract the problems, e. g. a reduction of the CFL number, relaxation procedures or smoothing 

functions. 

IV. Measurement campaigns and test matrix 

Within the VITAL project, several acoustic measurement campaigns have been conducted at the test 

facility described above. They can be subdivided into two categories, the first involving geometric variations 

of the axial gapping between the subsequent blade rows, and the second acoustic modifications of the 

TEC to reduce the noise generation or downstream radiation. The following listing summarizes the three 

configurations dealt with in this paper: 

• a reference or datum configuration (REF), 

• a maximum distance configuration (MAX), 

• a minimum distance configuration (MIN). 



For an easy modification of the axial gappings 

between the stator and the rotor, as well as the 

rotor and the EGV, the rig uses a modular design 

involving several axial spacer rings which can be 

inserted or removed at the respective locations. 

They are marked green in figures 6 and 7 showing 

longitudinal views of the turbomachinery section 

of the rig for the three axial gapping variations. 

Keeping the overall length constant, the MAX and 

MIN configurations are set by changing the spacer 

rings adjacent to the rotor. This corresponds to 

a change of the axial gapping-to-chord ratio by 

±45% and ±41.5% for the two locations statorrotor 

and rotor-EGV, respectively. It has to be 

noted at this point, that both distance variations 

involve simultaneous modifications of the gapping 

both ahead of and behind the rotor, i. e. reduced 

or increased distances to the up- and downstream 

airfoils. Separate variations or combinations of both have not been tested. 

Figure 6. reference or datum configuration 

(a) MAX configuration (b) MIN configuration 

Figure 7. Distance variations in STTF rig 

Concerning the operating conditions, all configurations have been computed and tested at conditions 

corresponding to the acoustic design points approach, cutback, and sideline (as described above). With 

respect to the experimental studies, this involves the exact adjustment of the flow parameters and operating 

conditions, and the recording of the pressure time series at all measurement positions by the traversing 

microphone array. The analysis of the data has then been carried out using the procedure described above. 

On the numerical side, a multitude of computations had to be performed to approximate each of the 

configurations. This relates to the linearization of the equations and the separation of the perturbations 

from the noise generation processes. Accordingly, for each blade ↔ vane interaction, each perturbation 

generation mechanism, and each harmonic of the perturbation, a separate computation had to be started. 

For illustration, table 3 summarizes the combinations of these parameters. 

Starting at the top, 3 interactions of relatively moving blade rows can be identified from the setup of 

the 1 1 

2 stage LPT depicted in figure 2: directly obvious the interaction of stator ↔ rotor as well as rotor ↔ 

EGV. In addition, the wake of the IGV also interacts with the rotor after propagation through the stator 

cascade. However, this effect can be considered to be of lower magnitude than the direct interaction of the 

stator and the rotor as will be seen in the results. For the potential field interaction due to the relatively 

moving upstream or downstream stationary pressure field, only the direct interactions S1 ↔ B1 and B1 ↔ 

EGV are taken into account. 



mechanisms 

Table 3. Numerical test matrix 

interactions IGV → B1 S1 → B1 B1 → EGV 

wake x x x 

potential ↓ x x 

potential ↑ x x 

disturbance harmonic 1 & 2 sound harmonic 1 & 2 BPF 

With respect to the harmonic order of the disturbance and the generated sound field, the analysis has 

been limited to the first and second orders. a The combination of these parameters yields 14 computations 

per operating condition and accordingly 42 computations per configuration. It should be mentioned again 

at this point that all computations are coupled solutions at least from the origin of noise generation to the 

downstream end of the test rig to compare with the experimental results. This includes scattering effects at 

downstream blade rows with respect to azimuthal mode order and frequency in the general case. b 

V. Comparison of experimental and numerical results 

In this section, selected results of the experimental and numerical data of the several configurations 

(described in the previous section) will be compared and analyzed. They focus both on the correlation 

between measurements and simulations, and on the acoustic effects of the modifications to the geometry 

and/ or the TEC design. 

V.A. Reference or datum configuration 

At first, a comparison of the results of the experimental investigations and the two numerical/ analytical 

studies will be given based on the reference configuration. For this purpose, the modal results at the LPT 

exit have been summed up, in the numerical study involving also a superposition of the three characteristic 

disturbance generation mechanisms. Figure 8 depicts the corresponding PWLs at 1 BPF. 

Figure 8. Comparison of PWL at 1 BPF; datum configuration; EXP ↔ SIM 

aThis is related to the fact that the 1 & 2 BPF can be accurately resolved at all operating conditions with the given 

microphone setup and sampling rate. 

bHowever, in this specific setup, no frequency scattering can be observed due to the limitation to only one rotor. 



It can be seen that the PWLs of the three studies agree quite well at all three OPs, especially the 

Linearized Euler (Lin3d) calculations and the measurements (DLR) which are within 1.5 dB at cutback and 

exactly identical at sideline. The TCLOW results agree very well at approach but slightly underestimate 

the noise at cutback and sideline. In contrast, with respect to the Lin3d results there is a slight variation to 

a lower PWL at approach. This can be related to the fact that the LPT of the STTF rig involves a cut-off 

design of the stator ↔ rotor interaction at this operating condition. However, as the EGV is not too far 

apart axially, and due to the convective effect of the flow, at least part of the acoustic energy is scattered at 

the EGV into cut-on modes, i. e. the cut-off mode m=+24 is scattered into the cut-on modes m=+9, and 

m=-6. These mode orders then contribute to the measured sound field at the LPT exit and can be still 

identified in the modal spectrum as depicted in figure 9. In this representation, the colored arrows indicate 

the interaction source of the modes attributed to the three blade-vane interactions. 

Figure 9. Modal PWLs at approach; 1 BPF; datum configuration; EXP ↔ SIM 

In comparison to the experimental results, for which all cut-on mode orders are represented in the 

modal spectrum, the numerical results yield only distinct peaks at the mode orders corresponding to direct 

combinations of the interacting blade rows. c As can be seen, at approach these relevant mode orders are not 

clearly (or at all) separated from the ’background noise’. This fact clearly complicates the exact analysis of 

the modal PWL from the experimental results and thus constitutes the largest inaccuracy. In addition to 

that, two mode orders are predominant that cannot be directly explained (i. e. modes m=-5 and m=+10). 

However, comparing the Lin3d and measurement results, the deviations are consistently lower for the IGV 

→ B1 and B1 ↔ EGV interactions (which are both cut-on) than for the S1 ↔ B1 interaction exciting the 

cut-off mode m=+24, which can only be detected by its scattered components. 

Articulately different at cutback and sideline where the relevant azimuthal mode numbers are clearly 

separated from the background noise, especially the (now cut-on) modes due to the stator ↔ rotor interaction 

(compare results at sideline, 1 BPF in figure 10). For a better comparison between the measurements and 

predictions, figure 11 just concentrates on the relevant modes at the operating points cutback and sideline. 

On an average basis, the results agree very well for these conditions where all primary interactions are cut-on 

at source. 

c It has to be noted that for the PWL comparison in figure 8 only these ’relevant’ modes have been taken into account. 



Figure 10. Modal PWLs at sideline; 1 BPF; datum configuration; EXP ↔ SIM 

At the second harmonic of the BPF (2 BPF), the correlation of the results is slightly less pronounced 

as can be seen in figure 12. As above, the largest deviations can be observed at the OP approach due to 

the cut-off modes of the (otherwise) dominant interaction S1 ↔ B1. At sideline, all three results are within 

1.5 dB, whereas at cutback the experimental exhibits an even higher PWL than at sideline, a result which 

cannot really be explained at this stage. Related to this, it has to be mentioned that the accuracy of the 

analysis process approaches the tolerable limit at 2 BPF for these higher power OPs. 

With respect to the source generation origins, i. e. the interactions generating the acoustic modes detected 

at the LPT exit, the results confirm the trend mentioned above (at the analysis of the modal spectra): 

As depicted in figure 13(a), at approach all three interactions yield similar PWLs with a slight dominance 

of the IGV ↔ rotor interaction at 1 BPF. At the second sound harmonic, the last (most downstream 

located) interaction contributes the most to the overall PWL. In contrast, at the higher power settings of 

the LPT (here: at cutback), the stator ↔ rotor interaction mode (m=+24) becomes cut-on and in parallel 

the dominating noise source, both at 1 and 2 BPF (compare figure 13(b)). 

(a) cutback (b) sideline 

Figure 11. Comparison of relevant modal PWLs at cutback and sideline; 1 BPF; datum configuration; EXP ↔ SIM 



Figure 12. Comparison of PWL at 2 BPF; datum configuration; EXP ↔ SIM 

In addition, in this plot it becomes even more obvious that especially the second harmonic of the IGV 

→ B1 interaction is recorded at a strongly reduced level. This can be related both to the noise generation 

mechanism responsible for the propagating sound waves as well as to the distance effect. In this case only 

the interaction of the IGV wake convected through the stator row with the rotor is taken into account, no 

potential field interaction contributes to the noise generated by these two blade rows. As can be expected, 

this results in a considerably lower PWL of this interaction compared to the other ones, especially the stator 

↔ rotor interaction. 

In the next section, the effect of the distance variation of the rotor-neighboring stages will be analyzed 

and compared between predictions and measurements. However, it has to be noted that relative to the pretest 

predictions, deviations of the aerodynamic conditions have been observed in the measurements probably 

due to additional losses in the rig and deviations from e. g. the idealized inflow conditions. Accordingly, all 

shown results of the two numerical tools have been based on post-test predictions after an adjustment of the 

aerodynamic parameters, i. e. total pressure at the inlet to the turbomachinery section of the rig, mass flow, 

total pressure profiles at the three measurement planes (compare figure 4). 

(a) approach (b) cutback 

Figure 13. Comparison of interaction sources at approach and cutback; datum configuration; SIM 



In contrast, due to time restrictions, it was not possible to adapt the Lin3d results of the following 

measurement campaign to the actual measured conditions prior to the publication of this paper. Therefore, 

they have not been included in the next section but will be updated in the near future. 

V.B. Distance variations 

As shown above (in figures 6 and 7), two variations of the distances to the rotor have been tested at the 

STTF rig. The overall results of the PWL at the LPT exit for the three configurations MAX - datum - MIN 

are summarized in the plots in figure 14. d As before, the experimental results (DLR) are compared to the 

numerical studies, in this case only the analytical tool (MTU TCLOW). 

(a) cutback (b) sideline 

(c) approach 

Figure 14. Comparison of OAPWL; distance variations; EXP ↔ SIM 

The analytical and experimental results agree quite well at all three operating conditions with slightly 

larger deviations at sideline, MIN. However, the overall trend of increasing PWLs with decreasing distances 

can be clearly confirmed. In addition, the trend of an increasing gradient of the PWL curve with decreasing 

axial gappings (larger delta between datum ↔ MIN than MAX ↔ datum) becomes quite obvious. It is 

expected that also the Lin3d results will show a very similar behavior after a recomputation of all interactions 

and source mechanisms based on an updated aero solution (as mentioned above). 

d In this representation, the results of the first and second BPF have been summed up energetically. 



VI. Summary and conclusion 

Within the VITAL project, several acoustic measurement campaigns have been carried out on a cold flow 

1 1 

2 stage LPT rig at the Graz University of Technology. These have been compared to aeroacoustic predictions 

generated by two codes of differing fidelity, a semi-empirical/ analytical code and a 3d CAA code based on 

the LEE. It has been shown that especially the LEE results agree very well with the measurement data after 

an adjustment of the flow parameters to ensure consistent boundary conditions. However, in an operating 

condition where the dominating stator-rotor interaction is cut-off, larger deviations can be observed. These 

conclusions have been derived from a reference configurations as well as two distance variations involving 

the axial gappings of the stator, rotor, and the EGV. 

Acknowledgments 

This work has been carried out in the framework of the European research project VITAL (EnVIronmen- 

TALly Friendly Aero Engine) under grant AIP4-CT-2004-012271. The support of the European Union is 

gratefully acknowledged. 

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