**4. Research on TECs**

**Figures 3** and **4** show the sound power level (PWL) of different azimuthal mode orders m that are cuton. The PWL is evaluated for the first blade passing frequency (BPF) at the operating point approach. The black bars in both figures indicate the sound power level of sound waves propagating in flow direction *PW L* <sup>+</sup> and the gray bars show the PWL of propagating waves against flow direction *PW L* <sup>−</sup> . The dotted lines in the figures indicate the noise level of the measurements. The microphones are mounted on a vibrating machine which causes some systematic error.

**Figure 3.** Sound power level of the reference TEC (left) and the leaned TEC (right).

**Figure 4.** Sound power level of the inverse cutoff TEC (left) and the H‐TEC (right).

Together with noise generated due to the evaluation technique and noise related to the measurement error the depicted level of overall noise floor can be assumed. **Figure 3** (left and right) shows the sound power level for the reference TEC and the leaned TEC, respectively. For the leaned TEC it can be seen that the overall PWL is about 8 dB lower than that of the reference TEC. Comparison of the inverse cutoff TEC (**Figure 4** (left)) with the reference TEC shows that *PW L* <sup>±</sup> is about 5 dB lower for the inverse cutoff TEC. Figure 3 (left) and **Figure 4** (right) show again the sound power level of different azimuthal mode order m that are cuton for the two TEC configurations (reference and H‐TEC). Both are again evaluated for the first BPF at operating point approach. The sound power level of the H‐TEC configuration is about 1 dB lower than for the reference TEC. The noise reduction levels mentioned above also contain non‐representative modes that belong to interaction modes with certain test rig components. These modes are created for example due to interaction of the sound field and the rear supporting struts at rig exit resulting in additional scattered modes. The difference in overall sound power level of the first BPF *∆PWL* =*PWL* − *PW L ref* in and against flow direction as well as the sum of both is also given in **Table 4**.


**Table 4.** Overall sound power level change for 1BPF.

**4. Research on TECs**

14 Recent Progress in Some Aircraft Technologies

propagating in flow direction *PW L* <sup>+</sup>

against flow direction *PW L* <sup>−</sup>

systematic error.

shows that *PW L* <sup>±</sup>

**Figures 3** and **4** show the sound power level (PWL) of different azimuthal mode orders m that are cuton. The PWL is evaluated for the first blade passing frequency (BPF) at the operating point approach. The black bars in both figures indicate the sound power level of sound waves

measurements. The microphones are mounted on a vibrating machine which causes some

**Figure 3.** Sound power level of the reference TEC (left) and the leaned TEC (right).

**Figure 4.** Sound power level of the inverse cutoff TEC (left) and the H‐TEC (right).

Together with noise generated due to the evaluation technique and noise related to the measurement error the depicted level of overall noise floor can be assumed. **Figure 3** (left and right) shows the sound power level for the reference TEC and the leaned TEC, respectively. For the leaned TEC it can be seen that the overall PWL is about 8 dB lower than that of the reference TEC. Comparison of the inverse cutoff TEC (**Figure 4** (left)) with the reference TEC

(right) show again the sound power level of different azimuthal mode order m that are cuton for the two TEC configurations (reference and H‐TEC). Both are again evaluated for the first

is about 5 dB lower for the inverse cutoff TEC. Figure 3 (left) and **Figure 4**

and the gray bars show the PWL of propagating waves

. The dotted lines in the figures indicate the noise level of the

Evaluating only engine relevant specific airfoil interactions shows a different trend than that depicted in **Table 4**. **Table 5** shows the azimuthal airfoil interaction mode orders (cuton modes) and also scattered modes calculated according to the equation of Tyler and Sofrin [5]. Mode orders higherthan ±17 are cut off. The main interaction mode of the rotor with the TEGV of the reference TEC, leaned TEC, and the H‐TEC is cuton and clearly visible (see **Figures 3** and **4**).


**Table 5.** Airfoil interaction modes m.

**Table 6** shows how the sound power level changes (compared to the reference TEC) when considering only the main airfoil interaction modes given in **Table 5**. It can be seen that the inverse cutoff TEC achieves the highest reduction in sound power level (14 dB) of all investi‐ gated configurations. The leaned TEC reduces PWL by about 11 dB, but that means that it is still twice as "loud" as the inverse cutoff TEC. The aerodynamically optimized H‐TEC even increases the sound power level by about 2 dB. At aero design point both TECs, the refer‐ ence TEC and the leaned TEC have the same PWL. The aerodynamically optimized H‐TEC is about 1 dB louder than the reference TEC and the inverse cutoff TEC reduces PWL by about 3 dB which is half as loud as the reference TEC.


**Table 6.** Sound power level change for 1BPF (airfoil interaction modes).

Considering only modes due to the TEC interaction and the scattering of modes at the TEC it can be seen that the inverse cutoff TEC even reduces PWL by 94 dB(!), because that mode is cut off. The leaned TEC has a reduced PWL of 11 dB and the H‐TEC is again increasing the PWL by about 2 dB. In addition to the acoustic measurements five‐hole probe measure‐ ments have been performed in order to compare the aerodynamics (loss estimation) of the three different TEC configurations with that of the state‐of‐the‐art reference TEC.

**Figure 5.** Mach number distribution (left) and yaw angle distribution (right) at TEC inlet.

#### **4.1. TEGV inlet flow: plane C**

Comparing the TEGV inlet conditions of the different measurement campaigns for the different TEC configurations reveals only minor differences in flow quantities due to the upstream potential effect of the turbine exit guide vanes of TEC. The circumferentially mass averaged radial Mach number distribution show insignificant differences between the configurations (**Figure 5** (left); 1 tick mark=0.1). Also the yaw angle is almost the same for all three configurations as can be seen in **Figure 5** (right) (1 tick mark=20°), the difference is also negligible. The largest difference between the TEC configurations can be seen at the tip above 90% channel height.

#### **4.2. TEGV outlet flow: plane D0, D**

is about 1 dB louder than the reference TEC and the inverse cutoff TEC reduces PWL by

Considering only modes due to the TEC interaction and the scattering of modes at the TEC it can be seen that the inverse cutoff TEC even reduces PWL by 94 dB(!), because that mode is cut off. The leaned TEC has a reduced PWL of 11 dB and the H‐TEC is again increasing the PWL by about 2 dB. In addition to the acoustic measurements five‐hole probe measure‐ ments have been performed in order to compare the aerodynamics (loss estimation) of the

three different TEC configurations with that of the state‐of‐the‐art reference TEC.

**Figure 5.** Mach number distribution (left) and yaw angle distribution (right) at TEC inlet.

Comparing the TEGV inlet conditions of the different measurement campaigns for the different TEC configurations reveals only minor differences in flow quantities due to the upstream potential effect of the turbine exit guide vanes of TEC. The circumferentially mass averaged radial Mach number distribution show insignificant differences between the

**4.1. TEGV inlet flow: plane C**

**Configuration** *∆PW L* **<sup>+</sup>** *∆PW L* **<sup>−</sup>** *∆PW L* **<sup>±</sup>** Reference TEC 0.0 [dB] 0.0 [dB] 0.0 [dB] Leaned TEC -11.3 [dB] -9.4 [dB] -11.2 [dB] Inverse cutoff TEC -13.8 [dB] -17.7 [dB] -14.0 [dB] H-TEC +2.4 [dB] -11.5 [dB] +2.0 [dB]

about 3 dB which is half as loud as the reference TEC.

16 Recent Progress in Some Aircraft Technologies

**Table 6.** Sound power level change for 1BPF (airfoil interaction modes).

**Figure 6**. shows Mach number and yaw angle distribution downstream of the TEGVs. H‐TEC and inverse cutoff TEC show a much more uniform distribution of the yaw angle at TEGV exit than the leaned and the reference TEC due to the higher vane count, thus, significantly improving the inflow conditions to a following component such as a mixer. Due to the fact that the inverse cutoff and the H‐TEC does not consider additional blockage because of the higher TEGV count the Mach number in plane D could be possibly slightly higher than for the reference and the leaned TEC (same TEGV count).

**Figure 6.** Mach number distribution (left) and yaw angle distribution (right) at TEC outlet.

#### **4.3. Loss estimation**

A rough estimation of the total pressure loss coefficient *<sup>ζ</sup>* <sup>=</sup> *<sup>p</sup>*˜¯*<sup>t</sup>*,*<sup>C</sup>* <sup>−</sup> *<sup>p</sup>*˜¯*<sup>t</sup>*,*<sup>D</sup> <sup>p</sup>*˜¯*<sup>t</sup>*,*<sup>C</sup>* <sup>−</sup> *pex* from plane C upstream of the TEGVs to plane D downstream of the TEGVs is done. The total pressure has been mass averaged by means of the five‐hole probe data. **Table 7** compares the change of total pressure loss of the three configurations. The loss coefficients have been normalised with the total pressure loss coefficient *ζref* of the reference configuration. Inverse cutoff, the leaned TEC as well as the H‐TEC configuration show higher loss coefficients than the reference TEC. That means that all configurations produce higher losses for the acoustically important operating point approach and it seems that they are more sensitive at off design conditions. However, the difference between the leaned and the inverse cutoff TECs is very small and within the measurement uncertainty.


**Table 7.** Total pressure loss.

Further, it can be seen that the losses increase from plane D0 to D for the H‐TEC. The losses are nearly constant for the inverse cutoff TEC from plane D0 to D. However, pressure loss measurements at aero design point showed a significant loss reduction from plane C to D for both, the aerodynamically optimised H‐TEC and the inverse cutoff TEC when compared with the reference TEC <sup>ζ</sup> <sup>ζ</sup>ref <sup>=</sup>0.9. A comparison between the leaned TEC and the reference TEC show similar losses so that *<sup>ζ</sup> <sup>ζ</sup>ref* <sup>≈</sup>1.

**Figure 7.** Comparison of modal PWL at LPT exit for two different stages; identical shaft power.

During the last decades a lot of institutions investigated the possibility to reduce engine weight by reducing the blade count of a low pressure rotor leading to highly loaded or even ultra‐ high‐loaded turbine blades. Therefore, also the acoustical behaviour of such a rotor compared to a state‐of‐the‐art design is investigated. At first, the results for the designated, acoustically relevant off‐design point approach for the two rotors, called OP1 for the datum stage (higher rotational speed) and OP2 for the high loaded stage (lower rotational speed), are shown. Due to the change in loading of 20%, the resulting rotational speed for the two geometries differs by 10% (see **Table 1**). **Figure 7** shows a comparison of the sound power level of different azimuthal mode orders evaluated for the first blade passing frequency (BPF) for the two different rotor setups. The cutoff criteria for the two operating points shows that for operating point 1 (OP1), the number of propagating modes is 39 (±19) whereas for operating point 2 (OP2), the number of propagating modes is only 35 (±17). Positive modes rotate against the rotational direction of the rotor and negative modes rotate in the same direction as the rotor. The figure also clearly shows that modes dominating the sound field are different for the two rotor configurations.

**Figure 8.** Comparison of modal PWL for propagating modes (rotor‐TEGV‐interaction).

the difference between the leaned and the inverse cutoff TECs is very small and within the

Further, it can be seen that the losses increase from plane D0 to D for the H‐TEC. The losses are nearly constant for the inverse cutoff TEC from plane D0 to D. However, pressure loss measurements at aero design point showed a significant loss reduction from plane C to D for both, the aerodynamically optimised H‐TEC and the inverse cutoff TEC when compared with

Configuration C-D C-D0 Ref. TEC 1.0 - Leaned TEC ∼1.6 - Inverse cutoff TEC ∼1.6 ∼1.6 H-TEC ∼1.3 ∼0.9

**Figure 7.** Comparison of modal PWL at LPT exit for two different stages; identical shaft power.

During the last decades a lot of institutions investigated the possibility to reduce engine weight by reducing the blade count of a low pressure rotor leading to highly loaded or even ultra‐ high‐loaded turbine blades. Therefore, also the acoustical behaviour of such a rotor compared to a state‐of‐the‐art design is investigated. At first, the results for the designated, acoustically relevant off‐design point approach for the two rotors, called OP1 for the datum stage (higher rotational speed) and OP2 for the high loaded stage (lower rotational speed), are shown. Due to the change in loading of 20%, the resulting rotational speed for the two geometries differs by 10% (see **Table 1**). **Figure 7** shows a comparison of the sound power level of different

**Total pressure loss** *<sup>ζ</sup>*

*ζref*

<sup>ζ</sup>ref <sup>=</sup>0.9. A comparison between the leaned TEC and the reference TEC show

measurement uncertainty.

18 Recent Progress in Some Aircraft Technologies

**Table 7.** Total pressure loss.

the reference TEC <sup>ζ</sup>

similar losses so that *<sup>ζ</sup>*

*<sup>ζ</sup>ref* <sup>≈</sup>1.

**Figure 9.** Comparison of modal PWL for propagating modes (rotor‐stator‐interaction).

The dominant modes for the OP1 with the datum rotor are the modes m=‐6 and +10, towering the next highest modes by approximately 4 dB. While the mode ‐6 is predicted by Tyler and Sofrin, the mode m=+10 could not be calculated. When, due to the higher loading of the second stage, the rotational speed is lowered, the modal distribution of the resulting flow field changes. The dominant mode for OP2 is the mode ‐5, that is, 2 dB higher than the second highest mode m=+10. The overall sound power level of the two configurations is shown in **Figure 8**. It can be seen that the PWL of the high loaded rotor with reduced rotational speed is reduced by 0.7 dB. Considering only the propagating modes predicted according to the theory of Tyler and Sofrin, the high loaded rotor shows a 2.5 dB lower PWL. The difference in power level between propagating modes according to Tyler and Sofrin and considering all modes found in the measurements is approximately 5 dB. The highest modes found in the analysis (m=+10 for OP1 and m=‐5 for OP2) of the measurement data are not attributable to Tyler and Sofrin. It is assumed that non‐Tyler Sofrin modes are increased for the highly loaded stage. Also, the sound power due to airfoil lift is assumed to be higher for the high loaded rotor. In addition, **Figure 8** depicts the modal PWLs of the Tyler‐Sofrin modes of the interaction between the rotor and the TEGVs.

The modes according to Tyler and Sofrin are m=‐12, m=+3, and m=+18. For the datum stage the highest amplitude for the rotor‐TEGV interaction mode is for m=+3. For the high loaded stage both modes m=‐12 and m=+3 show the same power level amplitude. Mode m=‐18 is cut off for the high load stage. Its contribution to the overall power level for the datum stage is negligible. **Figure 9** depicts the modal sound power levels of the stator‐rotor interaction. The main interaction mode would be the mode m=+24 that is cut off.

However, due to the small axial distance between stator and rotor the mode has not fully decayed and a scattering of that mode at the TEGVs is possible resulting in additional modes m=+9 and m=‐6. The remaining interactions between the rotor and the inlet guide vane of the testrig are depicted in **Figure 10**.

**Figure 10.** Comparison of modal PWL for propagating modes (rotor‐IGV‐interaction).

The mode m=‐4 has the highest PWL for both configurations, but is still approximately 7 dB lower than the sum of the Tyler and Sofrin modes. The mode m=11 is significantly lower as well as the mode m=‐19. This mode is only cuton and therefore able to propagate for the datum stage but is cut off and cannot propagate for the high loaded one. Comparing the different PWLs due to the interactions according to Taylor and Sofrin, the largest contributors to the overall PWL for the high loaded stage are interactions between the rotor and the TEGV as well as the mode m=‐6, which is a rotor‐stator interaction. For the datum stage the rotor‐TEGV interaction (m=+3) and the rotor‐IGV interaction (m=‐4) are the main contributors to the overall sound power level.

#### **4.4. TEGV pressure loss estimation**

Also for that investigation a rough estimation of the total pressure loss coefficient from plane C upstream to plane D downstream of the TEGVs is done by means of the five‐hole probe data. **Table 8** shows the total pressure loss coefficients between both configurations. They have been normalised with the total pressure loss coefficient *ζref* of the reference configuration. The losses produced by the TEGV are slightly increased by about 16% for the high stage loading stage upstream the TEGV than for the datum stage. It seems that flow structures from the rotor are mixed out in that TEC region and that the high loaded rotor produces flow structures resulting in higher mixing losses.


**Table 8.** Total pressure loss.

Tyler and Sofrin. It is assumed that non‐Tyler Sofrin modes are increased for the highly loaded stage. Also, the sound power due to airfoil lift is assumed to be higher for the high loaded rotor. In addition, **Figure 8** depicts the modal PWLs of the Tyler‐Sofrin modes of the interaction

The modes according to Tyler and Sofrin are m=‐12, m=+3, and m=+18. For the datum stage the highest amplitude for the rotor‐TEGV interaction mode is for m=+3. For the high loaded stage both modes m=‐12 and m=+3 show the same power level amplitude. Mode m=‐18 is cut off for the high load stage. Its contribution to the overall power level for the datum stage is negligible. **Figure 9** depicts the modal sound power levels of the stator‐rotor interaction. The

However, due to the small axial distance between stator and rotor the mode has not fully decayed and a scattering of that mode at the TEGVs is possible resulting in additional modes m=+9 and m=‐6. The remaining interactions between the rotor and the inlet guide vane of the

The mode m=‐4 has the highest PWL for both configurations, but is still approximately 7 dB lower than the sum of the Tyler and Sofrin modes. The mode m=11 is significantly lower as well as the mode m=‐19. This mode is only cuton and therefore able to propagate for the datum stage but is cut off and cannot propagate for the high loaded one. Comparing the different PWLs due to the interactions according to Taylor and Sofrin, the largest contributors to the overall PWL for the high loaded stage are interactions between the rotor and the TEGV as well as the mode m=‐6, which is a rotor‐stator interaction. For the datum stage the rotor‐TEGV interaction (m=+3) and the rotor‐IGV interaction (m=‐4) are the main contributors to the overall

Also for that investigation a rough estimation of the total pressure loss coefficient from plane C upstream to plane D downstream of the TEGVs is done by means of the five‐hole probe data. **Table 8** shows the total pressure loss coefficients between both configurations. They have been normalised with the total pressure loss coefficient *ζref* of the reference configuration. The losses produced by the TEGV are slightly increased by about 16% for the high stage loading stage

main interaction mode would be the mode m=+24 that is cut off.

**Figure 10.** Comparison of modal PWL for propagating modes (rotor‐IGV‐interaction).

between the rotor and the TEGVs.

20 Recent Progress in Some Aircraft Technologies

testrig are depicted in **Figure 10**.

sound power level.

**4.4. TEGV pressure loss estimation**
