Active flow control by a novel combinational active thermal protection for hypersonic vehicles

Journal Pre-proof Active flow control by a novel combinational active thermal protection for hypersonic vehicles Jie Huang, Wei-Xing Yao PII:

S0094-5765(20)30053-9

DOI:

https://doi.org/10.1016/j.actaastro.2020.01.033

Reference:

AA 7857

To appear in:

Acta Astronautica

Received Date: 8 December 2019 Revised Date:

18 January 2020

Accepted Date: 24 January 2020

Please cite this article as: J. Huang, W.-X. Yao, Active flow control by a novel combinational active thermal protection for hypersonic vehicles, Acta Astronautica, https://doi.org/10.1016/ j.actaastro.2020.01.033. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2020 Published by Elsevier Ltd on behalf of IAA.

Active flow control by a novel combinational active thermal protection for hypersonic vehicles Jie Huanga∗, Wei-Xing Yaoa,b a

State Key Laboratory of Mechanics and Control of Mechanical Structures, Nanjing University of Aeronautics and Astronautics, Nanjing 210016, China b

Key Laboratory of Fundamental Science for National Defense-Advanced Design Technology of Flight Vehicle, Nanjing University of Aeronautics and Astronautics, Nanjing 210016, China Abstract: This paper studies the mechanism of combinational active thermal protection system (CA-TPS) of the

hypersonic vehicles by numerical method. Unlike the traditional configuration, the nozzle is fixed on the spike, and the directions of lateral jet and hypersonic free stream are perpendicular to each other. The CA-TPS has significantly weaker reattachment shock wave than traditional spiked blunt body. Furthermore, the low temperature gas ejected from the nozzle flows downstream to reduce the aeroheating directly. For the blunt body, the peak and total heat fluxes of CA-TPS are 37.34% and 23.97% lower than that of traditional spiked blunt body, respectively. In addition, this paper conducts the analysis of influencing factors of the CA-TPS. As the total-pressure-ratio of nozzle and the length of spike increase, the reattachment shock wave intensity gradually weakens, and more low temperature gas can be ejected to increase the thermal protection efficiency of the CA-TPS. As the nozzle moves downstream, the reattachment shock wave intensity gradually increases, thus reducing the thermal protection efficiency. The effects of the above influencing factors on the flow field and change rate of thermal protection efficiency of CA-TPS are also studied in this paper. Keywords: Hypersonic; Thermal protection; Blunt body; Lateral jet; Spike.

Nomenclature d, L

diameter and length of spike

L1

distance between stagnation point and nozzle.

L2

size of nozzle

m

mass flow rate of nozzle

∗

Corresponding author. E-mail address: [email protected] (J. Huang).

Maj

Mach number of pressure inlet

PR

total pressure ratio

P∞

static pressure

P0j

total pressure of lateral jet

Qmax

peak heat flux

Qw

wall heat flux

T∞

static temperature

Tw

wall temperature

T0j

total temperature of lateral jet

α

angle of attack

∆Qmax

percentage decrease of peak heat flux

∆Qt

percentage decrease of total heat flux

Subscripts max

maximum

j

jet

t

total

w

wall

1. Introduction The aeroheating is one of the main characteristics of hypersonic vehicles, the high heat flux causes the sharp rise of structural temperature [1, 2]. To reduce the temperature of the load-bearing structure such as the skin, the thermal protection system (TPS) needs to be designed on the surface of vehicles. The tile is the most widely used passive TPS [3–5]. As the Mach number of modern hypersonic vehicles increases rapidly, the wall heat flux rises sharply, and the vehicles will be in the extremely thermal environment. In such thermal environment, the thickness of the tile must be increased to ensure thermal protection performance, which will result in the overweight of the tile. In addition, the maximum temperature of the tile will also rise rapidly, and it is very likely to exceed the allowable temperature of the tile and to cause the tile damage. The limitations of the tile thickness and the temperature resistance of materials make it impossible to adapt to the extremely thermal environment for modern hypersonic

vehicles, so the new TPS must be found. The active TPS can overcome the above two shortcomings of the tile. It mainly contains two basic configurations, the spike and jet configurations. The thermal protection mechanism of the active TPS has important research value. The spike is the uniform cross-section rod fixed on the nose cone of vehicles to reduce the aeroheating. The investigations on its thermal protection mechanism have been conducted through experimental and numerical methods since the 1950s [6–10]. The free stream is compressed by the spike to form a bow shock wave. Although the velocity of free stream reduces after the compression of the spike, it is still supersonic flow, which can be compressed by blunt body to form a reattachment shock wave. From the perspective of whole flow characteristics, the spike destroys the original bow shock and transform it into a weaker oblique shock wave to achieve the purpose of reducing the intensity of shock wave and aeroheating. The reattachment point and peak heat flux are located at the waist of the nose cone. In addition, the installation of the aerodisk at the stagnation point of the spike can increase the thermal protection performance. The aerodisk increases the compression of hypersonic free stream, which reduces the velocity of the gas reaching the nose cone, thus reducing the reattachment shock wave intensity [11, 12]. Huang et al. [13] analyzed the thermal protection mechanism of the spike under large angle of attack and found that there is an installation angle for the spike that can minimize the peak heat flux of the nose cone, which is about 1.5° larger than the angle of attack. However, the spike configuration also faces many challenges. Due to the small radius of the spike, the spike will be subjected to huge aeroheating in hypersonic flows, the structural temperature of the spike will be very high to damage the spike [14–16]. In addition, the installation of the spike destroys the aerodynamic shape of the original vehicle. When the angle of attack of hypersonic vehicles is not zero, the spike will generate additional pitch and yaw moments, which will influence the control of the vehicle. The jet is another type of active TPS, the investigations on its thermal protection mechanism have been conducted since the 1960s [17–19]. As with the spike, the jet can compress the free stream in advance and reduce the velocity of the gas reaching the nose cone, thus reducing the reattachment shock wave intensity. Furthermore, the low temperature gas is ejected from the nozzle and can cool the wall directly [20–22]. The reattachment point and peak heat flux are located at the waist of nose cone. Compared with the spike, the jet does not have ablative problems and will not destroy the aerodynamic shape of the original vehicle, so it is a potential active TPS. Huang et al. [23] studied the effect of jet mode on aeroheating of the nose cone by CFD numerical method. The jet appears as the long and short penetration modes in hypersonic flows. The critical total pressure of the nozzle determines the

penetration mode. Improving total pressure under different modes can reduce the aeroheating nose cone, and the thermal protection efficiency is also increased when the jet undergoes the mode transition. In recent years, the scholars began to design the combinational configurations and studied their thermal protection mechanism, including the combination of the spike and front jet [24–27] and the combination of the spike and rear jet [28, 29]. The results show that the combinational configurations have higher thermal protection efficiency than single configuration, and the flow characteristics of the combinational configuration are also much more complex than single configuration. Although the combinational configuration has excellent thermal protection performance, most of the combinational configurations include the spike, which inherits some shortcomings of the spike. According to the previous investigations, the combinational active TPS with the spike and front jet has good thermal protection performance. However, this combinational configuration has some disadvantages. Since the front jet is fixed at the stagnation point of the spike, the direction of the front jet is opposite to that of hypersonic free stream. Only the total pressure of the nozzle is greater than the static pressure of stagnation point of the spike, can the gas be ejected. This paper considers a novel combinational active thermal protection system (CA-TPS). The basic components include the blunt body, the spike and the lateral jet. The directions of lateral jet and hypersonic free stream are perpendicular to each other, so the static pressure of stagnation point of the spike has no influence on the ejecting of the gas. The lateral jet can flow tangentially to the downstream under the influence of hypersonic free stream, which overcomes the disadvantage of the front jet. This paper adopts the numerical method to study the mechanism of the CA-TPS and conducts the analysis of influencing factors.

2. Geometric and numerical models The CA-TPS studied in this paper is shown in Fig. 1(c), it achieves the thermal protection by the active flow control. The blunt body is the combination of the hemisphere and cylinder, the spike is fixed at the stagnation point of the blunt body, and nozzle is fixed on the spike, the gas can be ejected from the nozzle equably. In order to validate the thermal protection efficiency of the CA-TPS, two configurations in Fig. 1(a) and Fig.1 (b) are also considered. In addition, the sizes of the three configurations in Fig. 1 is shown in Table1.

L3

L3 L D

D Spike Blunt Body

(a) Single blunt body

(b) Spiked blunt body L3 L2

L1 L d D Lateral Jet

(c) Spiked blunt body with lateral jet Fig. 1. Geometric models. Table 1 Sizes of three configurations. L2 (mm) L (mm) L1 (mm) 75 5 1.25

L3 (mm) 15

D (mm) 50

d (mm) 5

This paper defines the temperature of all walls as 300K. Both the hypersonic free stream and the jet gas are considered as the calorically perfect gas. According to the geometric models, the three-dimensional computational grid of the CA-TPS are generated. Fig. 2 presents the topology of grid, the grid of symmetry plane, the grid of outlet and the local grid of the CA-TPS. The walls of the spike and blunt body are defined as the isothermal wall, the nozzle is defined as the pressure inlet, and the outer boundary is defined as the far-field and the supersonic outlet. Tables 2 and 3 present the parameters of the far-field condition and the pressure inlet, respectively. The height of first layer grid for all walls is defined as 1×10-6m (y+<1). The thermal protection performance of the CA-TPS is studied by the computational fluid dynamics method. The AUSM+ scheme [30] and the Menter’s SST k-ω twoequation turbulent model [31] are employed in the numerical simulation. The convergence tolerance of the total heat flux of all walls is defined as 10-3W. Table 2 Parameters of far-field condition. Parameter Quantity Ma∞ Mach number P∞ (Pa) Static pressure T∞ (K) Static temperature α (°) Angle of attack

Value 5 21.96 247.02 0

Table 3 Parameters of pressure inlet. Parameter Quantity Ma0j Mach number

Value 1.5

P0j (Pa) T0j (K)

(a) Topology of grid

Total pressure Total temperature

1161.8 300

(b) Grid of symmetry plane

(c) Grid of outlet (d) Local grid Fig. 2. Computational grid of CA-TPS. This paper conducts the grid independence investigation of the numerical model. In this paper, three grid systems with the grid number of 1925463, 2515831 and 3138819 are established for the CA-TPS without lateral jet. This paper only considers the aeroheating of the blunt body, so Fig. 3 only presents its heat flux distribution. The analysis results have little difference between the grid 2 (2515831) and grid 3 (3138819). Therefore, the computational grid with the grid number of 2515831 is chosen for the numerical calculation.

Fig. 3. Wall heat fluxes under different grid systems.

3. Code validation for aeroheating The code validation is required before studying the thermal protection performance of the CA-TPS. The experimental model of the straight biconic body is chosen for the aeroheating analysis in this paper. The computational grid and boundary conditions of the straight biconic body are shown in Fig. 4, and the numerical methods adopted are the same as that in Section 2 (AUSM+ scheme and Menter’s SST k-ω model). The height of first layer grid for the wall is defined as 1×10-6m, and the total grid number is about 500000. Table 4 presents the parameters of the far-field condition, the wall temperature of the straight biconic body is 300K, and the geometric information can be referred to Ref. [32]. Fig. 5 presents the experimental and calculated results. The heat flux distributions Qw obtained by the experiment and calculation match well, and the relative error between experimental and calculated stagnation heat fluxes is only 3.1%. According to the analysis results, the CFD code adopted in this paper meets the accuracy requirement for the hypersonic aeroheating analysis.

Fig. 4. Numerical model of straight biconic body. Table 4 Parameters of far-field condition.

Parameter Ma∞ P∞ (Pa) T∞ (K) α (°)

Quantity Mach number Static pressure Static temperature Angle of attack

Value 9.86 59.92 48.88 0

Fig. 5. Experimental and calculated results.

4. Results and discussion 4.1. Comparison of flow field This paper only studies the aeroheating of the blunt body. The Mach number distributions and streamlines of the single blunt body, traditional spiked blunt body and CA-TPS are obtained by the numerical calculations (Fig. 6). According to Fig. 6(a) and Fig. 6(b), the free stream is compressed by the blunt body and the spike to form a bow shock wave. Although the velocity of the air reduces after the compression of the spike, it is still supersonic flow, which can be compressed by the blunt body to form another shock wave (reattachment shock wave shown in Fig. 6(b)). The aeroheating is mainly determined by the intensity of the shock wave. According to the flow fields, the reattachment shock wave shown in Fig. 6(b) is weaker than the bow shock wave shown in Fig. 6(a). Therefore, installing spike can achieve the purpose of reducing the aeroheating. In addition, a recirculation zone is formed for the traditional spiked blunt body, and the outer edge of the recirculation zone is shear layer. The position of the reattachment point indicates the peak heat flux will be located at the waist of the hemisphere. The flow field of the CA-TPS is shown in Fig. 6(c). The CA-TPS has a larger recirculation zone than the traditional spiked blunt body, thus will affect the reattachment point of the blunt body significantly. The most important is that the reattachment shock wave shown in Fig. 6(c) is significantly weaker than that shown in Fig. 6(b), and the low temperature gas can flow downstream to cool blunt body directly. Therefore, the CA-TPS has

weaker aeroheating than the traditional spiked blunt body for the blunt body. According to the local flow filed of the lateral jet in Fig. 6(c), due to the flow separation induced by lateral jet, two small recirculation zones are also formed near the nozzle, and the ejected gas obviously deflects downstream under the influence of hypersonic free stream. The above numerical results qualitatively validates that the CA-TPS has higher thermal protection efficiency than traditional spiked blunt body from the perspective of flow characteristics.

Ma

Ma

Bow Shock

Shear Layer Recirculation Zone Reattachment Shock

Bow Shock

(a) Single blunt body

(b) Spiked blunt body

Ma

Bow Shock

Bow Shock

Shear Layer Recirculation Zone 1

Recirculation Zone 2 Recirculation Zone 3

Reattachment Shock

(c) CA-TPS Fig. 6. Comparison of flow field. 4.2. Comparison of thermal protection efficiency The wall heat flux of the single blunt body, traditional spiked blunt body and CA-TPS are obtained by the numerical calculations (Fig. 7). In 0-30° region, the traditional spiked blunt body has significantly lower heat flux

than single blunt body, however it has higher heat flux in 30°-90° region. In 0-50° region, the CA-TPS has significantly lower heat flux than single blunt body, however it has slightly higher heat flux in 50°-90° region. In 090° region, the CA-TPS has lower heat flux than traditional spiked blunt body. The peak heat flux of the CA-TPS is located downstream of that of traditional spiked blunt body, which verifies qualitative conclusion in Section 4.1. In addition, the peak heat flux Qmax and total heat flux Qt are obtained. The total heat flux Qt is calculated by:

Qt = 2π R 2 ∫

θ = 90°

θ =θ 0

Qw sin θ dθ + 2π RL3

(1)

where R is the radius of the hemisphere, L3 is the length of the cylinder, and Qw is the wall heat flux. The peak heat flux Qmax of the CA-TPS is 49.23% and 37.34% lower than that of the single blunt body and traditional spiked blunt body, respectively. The total heat flux Qt of the CA-TPS is 24.60% and 23.97% lower than that of the single blunt body and traditional spiked blunt body, respectively. Although the peak heat flux Qmax of the traditional spiked blunt body is 19.06% lower than Qmax of single blunt body, corresponding total heat flux Qt is only 0.83% lower than Qt of single blunt body. It is because, in 30°-90° region, the traditional spiked blunt body has higher heat flux than single blunt body. Above calculated results in Fig. 7 and Table 5 quantitatively validate that the CA-TPS has higher thermal protection efficiency than the traditional spiked blunt body.

Fig. 7. Comparison of wall heat flux. Table 5 Comparisons of peak and total heat fluxes. Configuration Single Blunt Body Spiked Blunt Body Qmax (kW/m2) 76.07 61.57 Qt (kW) 0.1451 0.1439

4.3. Effect of total pressure of nozzle

CA-TPS 38.58 0.1094

Since the total pressure of the nozzle is adjustable, the investigation on the influence of total-pressure-ratio of the nozzle PR on thermal protection efficiency of the CA-TPS is conducted. The parameter PR is defined as follows:

PR =

P0 j

(2)

P0 ∞

where P0∞ and P0j are the total pressures of free stream and nozzle, respectively. The flow fields with the totalpressure-ratio PR being 0.1, 0.2, 0.3 and 0.4 are shown in Fig. 8, and the length of spike L is kept at 75mm. Fig. 9 presents the effect of parameter PR on two small recirculation zones 2 and 3 in front of and behind the nozzle. As the parameter PR increases, the bow shock wave moves further away from the wall, the intensity of the reattachment shock wave gradually weakens, and more low temperature gas can be ejected to increase the thermal protection efficiency. In addition, the recirculation zone 1 increases, and the positions of the reattachment point and peak heat flux gradually move downstream. Simultaneously, the flow separation induced by lateral jet gradually strengthens, the recirculation zones 2 and 3 gradually increases, and the recirculation zone 3 behind the nozzle gradually moves downstream.

Ma Ma

(a) PR=0.1

(b) PR=0.2

Ma

Ma

(c) PR=0.3 (d) PR=0.4 Fig. 8. Effect of parameter PR on flow field (L=75mm).

Fig. 9. Effect of parameter PR on recirculation regions 2 and 3 (L=75mm). The effect of parameter PR on wall heat flux distribution is obtained (Fig. 10). Table 6 presents the effect of parameter PR on peak heat flux Qmax, total heat flux Qt and corresponding percentage decreases. The percentage decreases ∆Qmax and ∆Qt are defined as follows:

Qmax , PR = 0 − Qmax  ×100% ∆Qmax = Qmax , PR =0   ∆Q = Qt , PR = 0 − Qt ×100%  t Qt , PR = 0 

(3)

where Qmax, PR =0 and Qt , PR = 0 are peak heat flux and total heat flux with parameter PR being 0, respectively. Increasing total-pressure-ratio of the nozzle can significantly reduce the wall heat flux distribution and increase the percentage decreases ∆Qmax and ∆Qt. When the length of the spike L is kept at 75mm, the percentage decreases ∆Qmax with the parameter PR of 0.1, 0.2, 0.3 and 0.4 are 37.34%, 57.41%, 69.95% and 78.09%, respectively, and the corresponding percentage decreases ∆Qt are 23.97%, 43.57%, 57.82% and 68.17%, respectively. Therefore, ∆Qmax is usually greater than ∆Qt under the same parameter PR and length of spike. As the parameter PR increases, the positions of the reattachment point and peak heat flux gradually move downstream, which verifies the above

qualitative analysis conclusion. Furthermore, the change rate of the wall heat flux distribution, ∆Qmax and ∆Qt gradually decreases. The effect of parameter PR on the mass flow rate of nozzle is obtained (Fig. 11). As the parameter PR increases, the mass flow rate of the nozzle increases significantly. There is a linear relationship between the mass flow rate and the parameter PR, and the slope in Fig. 11 is 4.519g/s. Therefore, the total-pressure-ratio PR can not increase indefinitely, the thermal protection efficiency and the mass flow rate of the nozzle should be considered comprehensively so as to select appropriate total-pressure-ratio.

(a) L=50mm

(b) L=75mm

(c) L=100mm (d) L=125mm Fig. 10. Effect of parameter PR on wall heat flux distribution. Table 6 Effect of parameter PR on peak and total heat fluxes. ∆Qmax (%) Qt (kW) ∆Qt (%) L (mm) PR Qmax (kW/m2) 0 70.34 0 0.1499 0 0.1 44.44 36.82 0.1187 20.81 50 0.2 29.79 57.65 0.0886 40.89 0.3 20.76 70.49 0.0659 56.04 0.4 15.06 78.59 0.0496 66.91 0 61.57 0 0.1439 0 75 0.1 38.58 37.34 0.1094 23.97 0.2 26.22 57.41 0.0812 43.57

0.3 0.4 0 0.1 0.2 0.3 0.4 0 0.1 0.2 0.3 0.4

100

125

18.50 13.49 56.13 35.47 24.05 17.09 12.65 53.16 34.27 23.30 16.62 12.49

69.95 78.09 0 36.81 57.15 69.55 77.46 0 35.53 56.17 68.74 76.50

0.0607 0.0458 0.1381 0.1034 0.0766 0.0574 0.0437 0.1338 0.1007 0.0749 0.0565 0.0437

57.82 68.17 0 25.13 44.53 58.44 68.36 0 24.74 44.02 57.77 67.34

Fig. 11. Effect of parameter PR on mass flow rate of nozzle. 4.4. Effect of length of spike The flow fields with the length of spike L being 50mm, 75mm, 100mm and 125mm are shown in Fig. 12, and the parameter PR is kept at 0.1. As the length of spike increases, the intensity of the reattachment shock wave gradually weakens. Therefore, increasing the length L can increase the thermal protection efficiency. The influence of length L on wall heat flux is obtained (Fig. 13). As the length of spike increases, the wall heat flux distribution gradually reduces, the recirculation zone 1 increases, the positions of the reattachment point and peak heat flux gradually move downstream. Furthermore, the change rate of wall heat flux distribution in Fig. 13 decreases. However, the length of spike has little effect on the small recirculation zones 2 and 3 induced by the lateral jet. According to Fig. 11 and Table 6, the length of spike has little effect on the mass flow rate of nozzle, and increasing length of spike significantly reduces the peak heat flux Qmax and total heat flux Qt. When the parameter PR is kept at 0.1, as the length of spike increases from 50mm to 125mm, the Qmax and Qt are reduced by 22.88% and 15.16%, respectively. In addition, the stiffness and strength of the spike will decrease at the same time. Therefore, the lengthdiameter ratio L/D can not increase indefinitely, and the thermal protection efficiency, stiffness and strength of the spike should be considered comprehensively so as to select appropriate length of the spike.

Ma Ma

(a) L=50mm

(b) L=75mm

Ma Ma

(c) L=100mm (d) L=125mm Fig. 12. Effect of parameter L on flow field (PR=0.1).

(a) PR=0.1

(b) PR=0.2

(c) PR=0.3 (d) PR=0.4 Fig. 13. Effect of parameter L on wall heat flux distribution. 4.5. Effect of position of lateral jet This paper studies the effect of position of the lateral jet on thermal protection efficiency of the CA-TPS. The parameter L1 in Fig. 1(c) describes the effect of position of the lateral jet, which is the distance between the stagnation point and the nozzle. The flow fields with the parameter L1 being 6mm, 27mm, 48mm and 69mm are shown in Fig. 14, the length of spike and parameter PR are kept at 75mm and 0.4, respectively. As the nozzle moves downstream, the reattachment shock wave increases gradually. Therefore, it can be concluded qualitatively that increasing the parameter L1 increases the aeroheating based on the flow fields in Fig. 14. Simultaneously, the nozzle is closer to the blunt body, which is good for the thermal protection. Therefore, it can be concluded qualitatively that increasing the parameter L1 reduces the aeroheating based on the gas cooling. According to the above discussions, in order to evaluate the effect of position of lateral jet on thermal protection efficiency of the CA-TPS, both the reattachment shock wave and the low temperature gas cooling should be considered. When the parameter L1 is 6mm, the recirculation zones 1 and 3 are far apart and have little interference with each other. When the parameter L1 is 27mm, the distance between recirculation zones 1 and 3 decreases, and they have large interference with each other. The recirculation zone 1 increases, the position of peak heat flux Qmax moves downstream gradually. When the parameter L1 is 48mm, the recirculation zones 1 and 3 have merged into a recirculation zone, the position of peak heat flux Qmax moves upstream gradually due to the decrease of recirculation zone 1. When the parameter L1 is 69mm, the jet structure forces the recirculation zone to exist only in the narrow space between the nozzle and blunt body, the position of peak heat flux Qmax continues moving upstream. Therefore, as the nozzle moves downstream, the recirculation zone 1 first increases and then decreases, and the position of peak

heat flux Qmax first moves downstream and then upstream. Finally, the recirculation zone 2 induced by lateral jet gradually increases. Recirculation Zone 3 Recirculation Zone 2

Recirculation Zone 1

Ma

(a) L1=6mm

Ma

Ma

(b) L1=27mm

Ma

(c) L1=48mm (d) L1=69mm Fig. 14. Effect of parameter L1 on flow field (L=75mm, PR=0.4). The effect of position of the lateral jet on wall heat flux is obtained (Fig. 15). The effect of position of the lateral jet on the peak heat flux Qmax, the total heat flux Qt and the mass flow rate of the nozzle m is shown in Table 7. As the nozzle moves downstream, the wall heat flux distribution gradually increases, causing thermal protection efficiency of the CA-TPS to decreases gradually. It indicates that the reattachment shock wave has greater influence than the low temperature gas on the thermal protection efficiency of the CA-TPS. The variation trend of position of the peak heat flux Qmax verifies the above qualitative analysis conclusion. As the parameter L1 increases from 6mm to 69mm, the Qmax and Qt are increased by 79.29% and 57.79%, respectively, and the parameter L1 has no effect on the mass flow rate m. Furthermore, the change rate of the wall heat flux rapidly increases, which is bad to the

thermal protection efficiency of the CA-TPS. Therefore, when designing the lateral jet, the nozzle should be installed at the front of the spike.

Fig. 15. Effect of parameter L1 on calculated results (L=75mm, PR=0.4). Table 7 Effect of parameter L1 on peak and total heat fluxes and mass flow rate of nozzle. L1 (mm) 6 27 48 69 Qmax (kW/m2) 13.57 15.55 18.07 24.33 Qt (kW) 0.0462 0.0529 0.0597 0.0729 m (g/s) 18.075 18.075 18.075 18.075

5. Conclusion This paper studies the mechanism of the CA-TPS by numerical method, the following conclusions are obtained: (1) The CA-TPS has significantly weaker reattachment shock wave than traditional spiked blunt body. In addition, the low temperature gas flows downstream to cool blunt body directly. For the blunt body, the peak heat flux of the CA-TPS is 49.23% and 37.34% lower than that of the single blunt body and traditional spiked blunt body, respectively. The total heat flux of the CA-TPS is 24.60% and 23.97% lower than that of the single blunt body and traditional spiked blunt body, respectively. The thermal protection mechanism of the CA-TPS is shock wave reduction and gas cooling. (2) As the total-pressure-ratio of nozzle and length of spike increase, the reattachment shock wave intensity gradually weakens, and more low temperature gas can be ejected to increase thermal protection efficiency of the CA-TPS. In addition, the change rate of thermal protection efficiency of CA-TPS gradually decreases simultaneously. As the nozzle moves downstream, the reattachment shock wave increases gradually, thus reducing

the thermal protection efficiency. Therefore, when designing the lateral jet, the nozzle should be installed at the front of the spike.

Acknowledgments This research was supported by the National Postdoctoral Program for Innovative Talents (BX20190152), Project funded by China Postdoctoral Science Foundation (2019M660118), Jiangsu Planned Projects for Postdoctoral Research Funds (2019K127) and Priority Academic Program Development of Jiangsu Higher Education Institutions.

Conflict of interest statement There are no conflicts of interest in relation to this manuscript.

References [1] M.G. Persova, Y.G. Soloveichik, V.K. Belov, et al., Modeling of aerodynamic heat flux and thermoelastic behavior of nose caps of hypersonic vehicles, Acta Astronauti. 136 (2017) 312–331. [2] G. Shoev, G. Oblapenko, O. Kunova, et al., Validation of vibration-dissociation coupling models in hypersonic non-equilibrium separated flows, Acta Astronaut. 144 (2018): 147–159. [3] J. Huang, P. Li, W. Yao, Thermal protection system gap analysis using a loosely coupled fluid-structural thermal numerical method, Acta Astronaut. 146 (2018) 368–377. [4] A.M. Oscar, S. Anurag, V.S. Bhavani, T.H. Raphael, L.B. Max, Thermal force and moment determination of an integrated thermal protection system, AIAA J. 48(1) (2010) pp. 119–128. [5] J. Huang, W. Yao, P. Li, Uncertainty dynamic theoretical analysis on ceramic thermal protection system using perturbation method, Acta Astronaut. 148 (2018) 41–47. [6] D.H. Crawford, Investigation of the flow over spiked-nose hemisphere-cylinder at Mach Number of 6.8, NACA 6(1) (1959) 112–118. [7] W.A. Mair, Experiments on separation of boundary layers on probes in front of blunt-nosed bodies in a supersonic air stream, Philos. Mag. 43(342) (1952) 695–716. [8] D.J. Maull, Hypersonic flow over axially symmetric spiked bodies, J. Fluid Mech. 8(04) (1960) 584–592. [9] C.J. Wood, Hypersonic flow over spiked cones, J. Fluid Mech. 12(04) (1962) 614–624.

[10] Y.A. Dem'yanov, V.N. Shmanenkov, On studying reverse flows in the zone of separation of turbulent boundary layer, J. Applied Math. Mech. 24(2) (1960) 237–239. [11] Q. Qin, J. Xu, S. Guo, Fluid-thermal analysis of aerodynamic heating over spiked blunt body configurations, Acta Astronaut. 132 (2017) 230–242. [12] W. Huang, L.Q. Li, L. Yan, et al., Drag and heat flux reduction mechanism of blunted cone with aerodisks, Acta Astronaut. 138 (2017) 168–175. [13] J. Huang, W. Yao, N. Qin, Heat reduction mechanism of hypersonic spiked blunt body with installation angle at large angle of attack, Acta Astronaut. 164 (2019) 268–276. [14] W. Huang, Z. Chen, L. Yan, et al., Drag and heat flux reduction mechanism induced by the spike and its combinations in supersonic flows: A review, Prog. Aerosp. Sci. 105 (2019) 31–39. [15] X. Sun, W. Huang, M. Ou, et al., A survey on numerical simulations of drag and heat reduction mechanism in supersonic/hypersonic flows, Chinese J. Aeronaut. 32(4) (2019) 771–784. [16] Z. Wang, X. Sun, W. Huang, et al., Experimental investigation on drag and heat flux reduction in supersonic/hypersonic flows: A survey, Acta Astronaut. 129 (2016) 95–110. [17] P.J. Finley, The flow of a jet from a body opposing a supersonic free stream, J. Fluid Mech. 26(02) (1966) 337– 368. [18] D.J. Romeo, J.R. Sterrett, Exploratory Investigation of the Effect of a Forward-facing Jet on the Bow Shock of a Blunt Body in a Mach Number 6 Free Stream, National Aeronautics and Space Administration, NASA TN D1605, 1963. [19] D.M. Bushnell, J.K. Huffman, Forward Penetration of Liquid Water and Liquid Nitrogen from the Orifice at the Stagnation Point of a Hemispherically Blunted Body in Hypersonic Flow, NASA TM X-1493, 1968. [20] K. Hayashi, S. Aso, Y. Tani, Experimental study on thermal protection system by opposing jet in supersonic flow, J. Spacecr. Rockets 43(1) (2006) 233–235. [21] W. Huang, Y.P. Jiang, L. Yan, et al., Heat flux reduction mechanism induced by a combined opposing jet and cavity concept in supersonic flows, Acta Astronaut. 121 (2016) 164–171. [22] W. Huang, R.R. Zhang, L. Yan, et al., Numerical experiment on the flow field properties of a blunted body with a counterflowing jet in supersonic flows, Acta Astronaut. 147 (2018) 231–240. [23] J. Huang, W. Yao, Z. Jiang, Penetration mode effect on thermal protection system by opposing jet, Acta

Astronaut. 160 (2019) 206–215. [24] J. Huang, W. Yao, X. Shan, Coupled fluid-thermal investigation on non-ablative thermal protection system with spiked body and opposing jet combined configuration, Chinese J. Aeronaut. 32(6) (2019) 1390–1402. [25] M. Ou, L. Yan, W. Huang, S. Li, L. Li, Detailed parametric investigations on drag and heat flux reduction induced by a combined spike and opposing jet concept in hypersonic flows, Int. J. Heat Mass Tran.126 (2018) 10–31. [26] J. Huang, W. Yao, A novel non-ablative thermal protection system with combined spike and opposing jet concept, Acta Astronaut. 159 (2019) 41–48. [27] W. Huang, A survey of drag and heat reduction in supersonic flows by a counterflowing jet and its combinations, J. Zhejiang Univ.-Sc. A 16(7) (2015) 551–561. [28] J. Huang, W. Yao, X. Shan, Numerical investigation on drag and heat reduction mechanism of combined spike and rear opposing jet configuration, Acta Astronaut. 155 (2019) 179–190. [29] M.B. Gerdroodbary, Numerical analysis on cooling performance of counterflowing jet over aerodisked blunt body, Shock Waves 24(5) 2014 537–543. [30] M.S. Liou, A sequel to AUSM: AUSM+, J. Comput. Phys. 129(2) (1996) 364–382. [31] F.R. Menter, Two-equation eddy-viscosity turbulence models for engineering applications, AIAA J. 32(8) (1994) 1598–1605. [32] C.G. Miller, Experimental and predicted heating distributions for biconics at incidence in air at Mach 10, NASA Technical Paper 2234, 1984.

1) The numerical calculation is conducted to study mechanism of CA-TPS. 2) Excellent thermal protection performance of CA-TPS is verified. 3) Effect of total pressure of nozzle is analyzed. 4) Effect of length of spike is analyzed. 5) Effect of position of nozzle is analyzed.

Declaration of Interest Statement Dear Editor, We would like to submit the enclosed manuscript entitled "Active flow control by a novel combinational active thermal protection for hypersonic vehicles", which we wish to be considered for publication in "Acta Astronautica". The authors state that there are no conflicts of interest in relation to this manuscript. We declare that we do not have any commercial or associative interest that represents a conflict of interest in connection with the work submitted. Jie Huang and Wei-Xing Yao.