Multi-revolution Transfers in Orbital Mechanics
Bruno Miguel Tomás Correia
Thesis to obtain the Master of Science Degree in Aerospace Engineering
Supervisors: Prof. Paulo Jorge Soares Gil Prof. Ron Noomen
Examination Committee
Chairperson: Prof. Filipe Szolnoky Ramos Pinto Cunha Supervisor: Prof. Paulo Jorge Soares Gil Member of the Committee: Prof. João Manuel Gonçalves de Sousa Oliveira
June 2016 ii To my family and friends.
”The important thing is to not stop questioning. Curiosity has its own reason for existing.” - Albert Einstein
iii iv Acknowledgments
This report is the result of an one-year project developed in two different countries, having started in The Netherlands while in Erasmus and finished here in Portugal. During this journey, some people helped in the actual topic of the thesis while others provided support even without having any knowledge about space engineering. Firstly, I am grateful to my two amazing supervisors who helped me a lot in both phases of this process. To professor Ron Noomen, who provided me this fascinating topic, I would like to thank for the support provided while studying in Delft University of Technology (TU Delft) even when he had no obligation to do so. To professor Paulo Gil, who accompanied me during the entire process while abroad and specially after returning to Instituto Superior Tecnico´ (IST), I would like to thank for accepting me as a MSc Thesis student and for all the assistance given and the ideas provided to help improving this work. Finally, I would like to thank to my family and friends for the help, support and understanding shown.
v vi Resumo
”Sera´ poss´ıvel melhorar trajectorias´ interplanetarias´ com a inclusao˜ de multi-revoluc¸oes˜ quando comparado com aquelas em que apenas revoluc¸oes˜ parciais sao˜ usadas?” Este e´ o mote para o desen- volvimento desta tese. Assim sendo, e´ feita uma comparac¸ao˜ entre aquelas trajectorias´ que por norma sao˜ utilizadas e as que incluem multi-revoluc¸oes˜ onde a sonda completa pelo menos uma revoluc¸ao˜ a` volta do Sol antes de interagir com um planeta. Uma poss´ıvel reduc¸ao˜ no delta-V necessario´ levara´ a melhorias nos racios´ entre a massa da payload e a massa total da sonda, podendo melhorar futuras missoes˜ espaciais. Metodos´ de optimizac¸ao˜ sao˜ usados para minimizar o delta-V necessario´ para uma transferenciaˆ interplanetaria´ envolvendo multiplos´ swing-bys, incluindo apenas manobras impulsivas quando a sonda interage com os planetas. O codigo´ desenvolvido para este fim usa as bibliotecas PyGMO e PyKEP como base de desenvolvimento. Tresˆ casos sao˜ estudados, comec¸ando por uma transferenciaˆ directa para Marte. Depois, uma missao˜ para Mercurio´ e´ estudada usando nao˜ so´ transferenciasˆ directas como tambem´ incluindo um swing-by em Venus.´ A ultima´ missao˜ tem como destino Jupiter,´ sendo primeiro estudadas trajectorias´ sem multi-revoluc¸oes˜ com diferentes sequenciasˆ planetarias´ a fim de obter a melhor delas, sobre a qual multi-revoluc¸oes˜ sao˜ incorporadas. Todos estes casos estudados levaram a melhorias do delta- V necessario´ com o uso das multi-revoluc¸oes,˜ especialmente a missao˜ para Jupiter´ onde melhorias tambem´ foram encontradas ao n´ıvel da durac¸ao˜ da viagem e da data de partida da Terra. Deste modo, a resposta a` pergunta apresentada inicialmente e´ afirmativa.
Palavras-chave: transferenciaˆ interplanetaria,´ multi-revoluc¸oes,˜ optimizac¸ao,˜ swing-by, delta-V.
vii viii Abstract
”Is it possible to improve the design of interplanetary trajectories by using multi-revolutions when comparing to the cases where only single-revolutions are used?” This is the question this thesis intends to answer by comparing solutions using just ”standard” single-revolution transfers with those including multi-revolutions, where the spacecraft makes at least one complete revolution around the Sun before interacting with a planet. If the delta-V budget reduces, then better payload to total mass ratios arise, which can improve future space missions. Optimization methods are used here to optimize the delta-V budget of an interplanetary transfer defined as a multiple gravity assist (MGA) problem, including only impulsive manoeuvres when close to the planets the spacecraft interacts with. To do this, PyGMO and PyKEP toolboxes are used as a base of the implemented code. Three cases are analysed here, starting with a direct transfer mission to Mars. Then, a mission to Mercury is studied using both direct transfers and others with a swing-by at Venus. The final case is that of a mission to Jupiter, being first analysed different planetary sequences having single-revolutions only; multi-revolutions are then included to the best solution obtained, a VEEGA trajectory. In all these cases better solutions were found in terms of delta-V budget by including multi-revolutions, specially the Jupiter mission where improvements were obtained not only in the delta-V (and mass) budget of the mission, but also in the duration of the transfer and departure epoch. So, the answer to the initial question is affirmative.
Keywords: interplanetary transfer, multi-revolutions, optimization, MGA problem, delta-V budget.
ix x Contents
Acknowledgments...... v Resumo...... vii Abstract...... ix List of Tables...... xiii List of Figures...... xv Nomenclature...... xvii
1 Introduction 1 1.1 Motivation...... 1 1.2 State of the Art...... 2 1.3 Thesis Outline...... 3
2 Fundamentals of Orbital Mechanics5 2.1 The Two-body Problem...... 5 2.1.1 Circular and Elliptical Orbits...... 6 2.1.2 Parabolic Orbits...... 7 2.1.3 Hyperbolic Orbits...... 7 2.2 Interplanetary Transfers...... 8 2.3 Orbital Manoeuvres...... 10 2.4 Gravity Assist Manoeuvres (Swing-bys)...... 12 2.4.1 Unpowered Gravity Assist...... 12 2.4.2 Powered Gravity Assist...... 14 2.5 Lambert’s ”Orbital Boundary-value Problem”...... 16 2.5.1 Izzo’s Final Form of Lambert’s Problem...... 17 2.5.2 Single-revolution versus Multi-revolutions...... 19 2.5.3 Izzo’s Approach to Solve Lambert’s Problem...... 20
3 The Multiple Gravity Assist Trajectory Problem 21 3.1 Problem Definition and Notation...... 21 3.2 General Form of the Problem and Simplifications...... 22 3.3 Final Form of the Problem...... 23
xi 4 Optimization 27 4.1 Optimization in the Design of Interplanetary Missions...... 27 4.2 Differential Evolution (Global)...... 28 4.2.1 Method Description...... 29 4.2.2 Variants of the Method...... 30 4.2.3 Control Parameters Selection...... 31 4.3 Simulated Annealing with Adaptive Neighbourhood (Global)...... 31 4.3.1 Method Description...... 32 4.3.2 Control Parameters Selection...... 33 4.4 Compass Search (Local)...... 33 4.4.1 Method Description...... 33
5 Implementation 35 5.1 Description of the MGA Problem Implementation...... 35 5.2 Description of the Optimization Implementation...... 38 5.3 Validation...... 41 5.3.1 Test I: MGA Problem...... 41 5.3.2 Test II: Optimization Process...... 42
6 Results and Discussion 45 6.1 Description of the Studied Cases...... 45 6.2 Interplanetary Transfer to Mars...... 46 6.2.1 Results for a Direct Transfer...... 46 6.2.2 Discussion...... 47 6.3 Interplanetary Transfer to Mercury...... 49 6.3.1 Results for a Direct Transfer...... 49 6.3.2 Results for a Transfer with Swing-by at Venus...... 50 6.3.3 Discussion...... 51 6.4 Interplanetary Transfer to Jupiter...... 54 6.4.1 Results for the Single-revolution Transfers...... 54 6.4.2 Results for a VEEGA Trajectory with Multi-revolutions...... 56 6.4.3 Discussion...... 57
7 Conclusions 61 7.1 Achievements and Final Remarks...... 61 7.2 Future Work...... 62
Bibliography 63
xii List of Tables
5.1 Planetary data used during the simulations...... 39 5.2 Control parameters for the optimization algorithms used in the archipelago...... 40 5.3 Comparison of the results for the Cassini1 problem for validation of the MGA class..... 42 5.4 Comparison of the results for the Cassini1 and Rosetta problems for validation of the optimization process...... 43
6.1 Boundaries for the optimization variables for a direct transfer from Earth to Mars...... 46 6.2 Results obtained for a direct transfer from Earth to Mars...... 47 6.3 Boundaries for the optimization variables for a direct transfer from Earth to Mercury.... 49 6.4 Results obtained for a direct transfer from Earth to Mercury...... 49 6.5 Boundaries for the optimization variables for a transfer from Earth to Mercury with a gravity assist at Venus...... 50 6.6 Results obtained for a transfer from Earth to Mercury with a gravity assist at Venus..... 50 6.7 Boundaries for the optimization variables for different single-revolution transfers from Earth to Jupiter...... 54 6.8 Boundaries for the optimization variables for a transfer from Earth to Jupiter using a VEEGA trajectory...... 56 6.9 Results obtained for a transfer from Earth to Jupiter using a VEEGA trajectory...... 57
xiii xiv List of Figures
2.1 Velocity vector diagram for an unpowered gravity assist...... 12 2.2 In-plane geometry for an unpowered swing-by as viewed in the planetocentric frame.... 13 2.3 Maximum change in specific energy of a spacecraft during a swing-by...... 14 2.4 In-plane geometry for a powered swing-by as viewed in the planetocentric frame...... 15 2.5 Example of an elliptical solution for Lambert’s problem with certain boundary conditions.. 16 2.6 Non-dimensional transfer time curves as a function of x for different values of λ and M.. 19
3.1 Sketch of the different phases in the multiple gravity assist (MGA) trajectory followed by the spacecraft...... 24
5.1 Structure of the MGA class containing the MGA Problem...... 37 5.2 Archipelago used for optimization and consisting of seven islands in rim topology...... 40
6.1 Detailed information of best solution obtained for a direct single-revolution transfer from Earth to Mars...... 48 6.2 Detailed information of best solution obtained for a direct transfer from Earth to Mars having one complete revolution...... 48 6.3 Detailed information of best solution obtained for a direct transfer from Earth to Mercury having two complete revolutions...... 51 6.4 Detailed information of best solution obtained for a transfer from Earth to Mercury with a swing-by at Venus with single-revolutions only...... 52 6.5 Detailed information of best solution obtained for a transfer from Earth to Mercury with a swing-by at Venus with one complete revolution in the first leg...... 53 6.6 Detailed information of best solution obtained for a transfer from Earth to Mercury with a swing-by at Venus having one complete revolution in both legs...... 53 6.7 Delta-V budget and duration for the best solutions for an interplanetary mission from Earth to Jupiter involving only single-revolutions...... 55 6.8 Detailed information of best solution obtained for a transfer from Earth to Jupiter using a VEEGA trajectory with single-revolutions only...... 58 6.9 Detailed information of best solution obtained for a transfer from Earth to Jupiter using a VEEGA trajectory with one complete revolution between both Earth swing-bys...... 60
xv xvi Nomenclature
Acronyms
ACT ESA’s Advanced Concepts Team
CS Compass Search
DE Differential Evolution
DSM Deep Space Manoeuvre
ESA European Space Agency
GA Gravity Assist manoeuvre
GTOP Global Trajectory Optimization Problems database
JPL NASA’s Jet Propulsion Laboratory
LEO Low Earth Orbit
MGA Multiple Gravity Assist problem
MGA-1DSM Multiple Gravity Assist problem with 1 Deep Space Manoeuvre per leg
MJD2000 Modified Julian Day starting on January 1, 2000
NASA National Aeronautics and Space Administration
PaGMO Parallel Global Multiobjective Optimizer
PyGMO Python Parallel Global Multiobjective Optimizer
SA Simulated Annealing
SA-AN Simulated Annealing with Adaptive Neighbourhood
SoI Sphere of Influence
TOF Time Of Flight
UTC Universal Time Coordinated
VEEGA Venus-Earth-Earth Gravity Assist trajectory
xvii Greek symbols
α Gravity assist: deflection angle [rad]
β Gravity assist: hyperbolic encounter angle [rad]
∆ Optimization: step range (CS) [-]
∆... Change in... [various]
γ Turning angle in a plane change manoeuvre [rad]
λ Lambert solver: non-dimensional parameter [-]
µ Gravitational constant of a body [m3/s2]
ψ Lambert solver: non-dimensional auxiliary variable [-]
θ True anomaly in an orbit [rad] Roman symbols
E Specific energy of a body [J/kg]
~i Unit vector [-]
~p Optimization: decision vector [various]
R~ Position vector of a planet [m]
~r Position vector of the orbiting body (spacecraft) [m]
T~ Thrusting force acting on the spacecraft [N]
V~ Velocity vector of a planet [m/s]
~v Velocity vector of the orbiting body (spacecraft) [m/s]
~v∞ Hyperbolic excess velocity [m/s]
~x State vector of the spacecraft [various] a Semi-major axis of the orbit [m]
B Gravity assist: impact parameter [m] c Lambert solver: Chord connecting the initial and final positions [m]
CR Optimization: crossover constant (DE) [-]
D Optimization: problem dimension [-]
E Eccentric anomaly in an elliptical orbit [rad] e Eccentricity of the orbit [-]
xviii F Optimization: scale factor (DE) [-] f Objective/cost function during optimization [various]
G Newtonian constant of gravitation [m3/(kg.s2)]
G Optimization: generation of the population [-] g Gravitational acceleration [m/s2]
H Hyperbolic anomaly in an hyperbolic orbit [rad]
Isp Specific impulse of the thrusters [s] lb Optimization: lower bound of a parameter [various]
M Lambert solver: number of complete revolutions around the Sun [-] m Mass of a body [kg]
Nε Optimization: number of temperature reduction cycles (SA-AN) [-]
NP Optimization: population size [-]
NS Optimization: number of same step vector parameter change cycles (SA-AN) [-]
NT Optimization: number of step vector change cycles (SA-AN) [-]
P Optimization: probability density function [-] p Semi-latus rectum of an orbit [m]
R Optimization: temperature reduction coefficient (SA-AN) [-] r Radius (distance) from the central body [m]
Rp Equatorial radius of a planet [m] rT Optimization: temperature reduction coefficient (CS) [-]
S Lambert solver: solution (x1 or x2) to use for multi-revolution case [-] s Lambert solver: semi-perimeter [m]
T Duration of an interplanetary leg [s]
T Lambert solver: non-dimensional transfer time [-]
T Optimization: temperature (SA-AN) [-] t Departure/encounter/arrival epoch [MJD2000/UTC] ub Optimization: upper bound of a parameter [various] x Lambert solver: non-dimensional iteration variable [-] y Lambert solver: non-dimensional auxiliary variable [-]
xix xx Chapter 1
Introduction
1.1 Motivation
In the last decades, mankind has witnessed the advent of space exploration. At this moment, all eight planets in the Solar System have been visited by probes from Earth and even Pluto was visited last year by the New Horizons probe [1]. One of the major challenges is to send heavier and more capable instruments to better analyse the different bodies orbiting the Sun. Multi-revolution transfers using high thrust propulsion, where the probe flies at least one complete revolution around the Sun between two planets, may provide better trajectory solutions in terms of the mass budget required for the interplanetary mission when comparing to a transfer with single-revolutions only. For that reason, the main question to be answered in this thesis is the following: is it possible to improve the design of an interplanetary trajectory by using multi-revolutions when comparing to the cases where only single-revolutions are used? This is one of several strategies to achieve the goal of a better ratio between payload mass and the total mass of the spacecraft, which translates into an increase of the payload mass that can be delivered to a certain planet or a decrease of the propellant mass needed to deliver a specific payload to its des- tination. However, this strategy may secondarily improve or deteriorate other mission’s characteristics such as the launch and/or departure windows and the overall duration of the interplanetary transfer. Single-revolution transfers are those in which the interplanetary trajectory leg connecting two bodies only completes a portion of a full revolution around the Sun, in other words, the initial position of the orbit is only acquired once, at the beginning [2]. On the other hand, multi-revolution transfers correspond to those cases where the initial position of the orbit is repeated at least once, meaning that the satellite makes one or more complete revolutions around the Sun in the same orbit before changing it by means of manoeuvres or encounters with other bodies such as intermediate planets. The reason for using multi-revolutions is the possibility of obtaining better geometry conditions for the trajectory, leading to a decrease in the delta-V and mass budgets. This is particularly relevant when employing intermediate swing-bys since more bodies are included in the trajectory design. Moreover, the introduction of these cases increases the degrees of freedom when designing the mission’s trajectory.
1 Besides those advantages, the use of multi-revolutions may bring a major drawback which is the considerable increase in the total time required to arrive at the target planet. Therefore, this scheme will only be considered when analysing missions to inner planets/bodies, up to the Asteroid Belt between Mars and Jupiter, or in the initial segments of a multiple gravity assist trajectory to the outer planets such as Jupiter and so forth, which still occur in that inner region.
1.2 State of the Art
Since the first successful interplanetary mission, Mariner 2, launched from Earth and arriving at Venus on December 14, 1962 [3], many other probes have left the Earth heading to different targets (planets and others) in the Solar System. The great majority of them used high thrust propulsion in order to escape the Earth and arrive at their destination. From all of those high thrust missions, most of them, if not all, only used single-revolution transfers, even when using multiple gravity assists such as the Galileo mission to Jupiter [4] or the Cassini-Huygens mission to Saturn [5]. Although knowledge of multi-revolution transfers exists for a long time, probably the same time as the knowledge of the single-revolution ones, they are not so commonly applied in missions, being merely treated as a theoretical result. The reason for this is the increase in the transfer time required and the difficulties introduced in the trajectory design process due to the increase in number of variables and cases to optimize. Nevertheless, it may be advantageous to study those trajectories where at least one of their legs consists of a multi-revolution transfer because of the possibility of decreasing the overall delta-V budget which translates in improvements in the payload to total mass ratio. Another interesting option is to use low thrust (electric) propulsion where gases are accelerated by means of electrical heating and/or by electromagnetic body forces which then accelerate the spacecraft
[6]. Here, the goal is attained mainly by the higher values of specific impulse Isp of those thrusters [7] instead of reducing the delta-V budget. This technique is becoming more and more attractive because of its advantages even if, just like in the high thrust multi-revolution case, the duration increases due to the low thrust provided [7]. Deep Space 1 launched in 1998 to asteroid 9969 Braille and comet Borrelly was the first successful deep space low thrust mission [8]. From that moment, other missions used low thrust as main propulsion system, such as the Dawn mission to the asteroids Ceres and Vesta [9]. Unlike chemical (high thrust) propulsion, electrical propulsion systems require much more power to operate, increasing the total power budget of the space probe and also correspond to much more complex systems. The electrical power can be obtained from solar energy whenever the spacecraft is at relatively small distances from the Sun (inner Solar System). However, for missions to the outer Solar System where the energy from the Sun is small, nuclear power sources are required: a nuclear reactor for heavy probes or a device heated by the decay of radioisotopes for lighter ones [10]. Another disadvantage of this particular strategy is that for huge changes in velocity, such as those required to inject the satellite into an interplanetary transfer (acceleration) or to inject it into an orbit around the target planet (deceleration), it is required thrusting forces which are, usually, beyond the capability of these low thrusters, even if they are more efficient in terms of propellant mass usage.
2 These two solutions addressed, multi-revolutions with high thrust propulsion and low thrust propul- sion, compete to the yet ”standard” high thrust single-revolution interplanetary trajectories. But, due to the requirements of better mass budgets in the future, they may see their importance increased, as it is already seen for the low thrust case with the increasing number of missions conducted. Lastly, although they both present the disadvantage of increasing the duration of the interplanetary flight, the multi-revolution solution uses the simpler high thrust systems and does not need to have secondary thrust systems due to the increase distances from the Sun for missions to the outer planets or the need to use nuclear propulsion, which still is a controversial topic for political and safety reasons [11]. This means that if better solutions using high thrust multi-revolution transfers are found when comparing to the single-revolution cases, maybe this strategy will start to be used more often in future space missions.
1.3 Thesis Outline
The structure of this work is presented now. Some fundamental concepts of orbital mechanics, which are the basis of every preliminary space mission design process, are studied in Chapter2. It is laid out in Chapter3 the actual problem to be optimized, using the different optimization methods presented in Chapter4. The implemented code is described and validated in Chapter5. Then, Chapter6 concerns with the results obtained for the different practical cases studied and respective discussion. Finally, Chapter7 contains the conclusions drawn regarding the topic of this thesis. By the end of reading this thesis the reader should be able to understand what multi-revolution transfers are and how they influence the overall design of an interplanetary trajectory. Further, it has answered the question presented in the second paragraph of Section 1.1 by analysing several cases of interplanetary missions departing from Earth and targeting three different planets in the Solar System: Mercury, Mars and Jupiter. The study of those cases is made by using a code to be developed in python and which optimizes interplanetary trajectories using high thrust. The PyGMO (Python Parallel Multiobjective Optimizer) [12] and PyKEP [13] toolboxes provided by the Advanced Concepts Team of the European Space Agency (ESA) are used as the basis for the development of that code since they already present functions and classes regarding some astrodynamics and optimization methods which facilitates its implementation.
3 4 Chapter 2
Fundamentals of Orbital Mechanics
In order to understand the different models used and to correctly analyse the results obtained regard- ing the different interplanetary transfers, it is essential to know some basic concepts of orbital mechanics and that is the goal of this chapter. It is divided into sections regarding different concepts of astrodynam- ics, starting from the simplest model of orbital motion up to the derivation of Lambert’s problem which plays an important role in the preliminary design of interplanetary trajectories.
2.1 The Two-body Problem
The simplest model for the motion of bodies in the Universe, such as the motion of planets around the Sun, is that of a two-body problem where the only force acting upon a certain body is the gravitational force of another body [7]. Under this model, the motion of body i due to the gravitational acceleration of body j in a non-rotating reference frame with origin in the latter is governed by the following equation:
d2~r m + m = −G j i ~r, (2.1) dt2 r3 where ~r is the position of body i with respect to body j, mi and mj are the masses of both bodies and G is the Newtonian constant of gravitation [14]. In the case of the motion of a spacecraft around a planet or the motion of any celestial body in the Solar System around the Sun, the mass of the orbiting body is much smaller when comparing to the mass of the body governing that motion, i.e., mi mj; therefore, the mass of the secondary body mi can be neglected in first approximation. Using the notation µ = Gmj, where µ is the gravitational constant of body j, the final equation for the acceleration of the orbiting secondary body is obtained in the reference frame considered [7]: d2~r µ = − ~r. (2.2) dt2 r3
This model, given by Equation (2.2), provides a closed solution (unlike models involving more than two bodies) whose orbit, flown by the secondary body, corresponds to a Keplerian orbit. This type of
5 orbits are defined by the following expression [7]:
p r = , (2.3) 1 + e cos θ being r the radius from the main body to the secondary one, p the semi-latus rectum, e the eccentricity of the orbit and θ the true anomaly, which is the angle between the radius and the direction of closest approach to the main body. The eccentricity of the orbit defines its geometry:
e = 0 : circle
0 < e < 1 : ellipse
e = 1 : parabola
e > 1 : hyperbola.
Circular orbits are a particular case of elliptic orbits where both the semi-major and the semi-minor axes are equal. So, they can be treated as a single case considering just that e < 1. Also, parabolic orbits are a limit case between elliptical and hyperbolic ones and they are inherently unstable since a small perturbation is sufficient to change it to one of the other two cases with e 6= 1 [7]. Further details about each type of Keplerian orbit are presented in Sections 2.1.1 through 2.1.3.
2.1.1 Circular and Elliptical Orbits
Circular and elliptical orbits are the only closed Keplerian orbits, which means that the orbiting body will continue to repeat that same orbit unless some perturbation changes its motion. Further, the specific energy of the body in that orbit is constant and negative (E < 0). If the semi-major axis of the orbit is denoted by a, with 0 < a < ∞, then the semi-latus rectum is given by [7]:
p = a(1 − e2) (2.4)
Also, the minimum and maximum distances from the orbiting body to the central body, also known as periapsis and apoapsis, are given by
rp = rmin = a(1 − e) (2.5a)
ra = rmax = a(1 + e). (2.5b)
Substituting p in Equation (2.3) by its expression in (2.4), it is possible to rewrite the equation of the orbit as a(1 − e2) r = , (2.6) 1 + e cos θ where it becomes clear that for a circular orbit, where the eccentricity is zero, the radius is always constant and equal to the semi-major axis. On the other hand, the difference between the periapsis and apoapsis distances increases as the eccentricity increases.
6 Another important result regarding the velocity of body i in its orbit is obtained from the Vis Viva Equation [7], v2 µ µ E = − = − , (2.7) 2 r 2a by mathematically manipulating it to obtain the following expression for the velocity v in an elliptical orbit:
s 2 1 v = µ − . (2.8) r a
For circular orbits, this velocity is always constant and simply given by
rµ v = , (2.9) c a meaning that the velocities at periapsis and apoapsis can be uniquely obtained from the local circular velocity vc and the eccentricity of the orbit [7]:
rµ 1 + e √ v = · = v 1 + e (2.10a) p a 1 − e c,p rµ 1 − e √ v = · = v 1 − e, (2.10b) a a 1 + e c,a corresponding them to the maximum and minimum velocities, respectively.
2.1.2 Parabolic Orbits
Even considering the approximation of a two-body problem and all its assumptions, these orbits are impossible to maintain due to their unstable nature. Nevertheless, they represent the limit from which the orbit is no longer closed and the orbiting body can escape the central body. When in a parabolic orbit, the orbital energy is equal to zero (E = 0) and the body comes from or arrives at an infinite distance with zero velocity with respect to the main body. In interplanetary missions, the most interesting result about these specific orbits is the velocity at periapsis since it represents the minimum velocity that a spacecraft has to possess in order to be able to leave the planet. This velocity, also known as escape velocity, is related to the local circular velocity and it is given by [7]: r2µ √ v = = 2v . (2.11) esc r c
2.1.3 Hyperbolic Orbits
Lastly, hyperbolic orbits are those whose eccentricity is greater than 1 and with positive specific energy (E > 0). In this orbit, a body departs from or arrives at an infinite distance still with a non-zero
finite velocity v∞ (corresponding to the minimum velocity in the trajectory) and has a maximum velocity at the periapsis. If the the semi-major axis is now considered to be negative instead of positive (a < 0), then all equations presented in Section 2.1.1 can still be used.
7 The magnitude of the velocity at an infinite distance from the central body, r → ∞, can be fully determined by the semi-major axis of the orbit [7]:
µ v2 = C = − . (2.12) ∞ 3 a
Moreover, the velocity at periapsis is also given by Equation (2.10a) and, for that reason, the ratios between vp and v∞ and between vp and the local circular velocity vc are respectively given by
v re + 1 p = (2.13) v∞ e − 1 and v √ p = e + 1 . (2.14) vc,p
One final remark about the instantaneous velocity of a certain body in an hyperbolic orbit is that it is always completely determined by the local escape velocity vesc and the velocity at infinity [7], being the following relation true for every point along the orbit:
2µ v2 = v2 + v2 = + C . (2.15) esc ∞ r 3
This relation is very important in the design of interplanetary missions since one usually knows the value of v∞ and by knowing the radius with respect to the central planet, its velocity can be easily de- termined. Furthermore, this relation is used to obtain the impulses required to provide to the spacecraft so that it enters a specific interplanetary trajectory, starting with a hyperbolic orbit leaving the departure planet (assuming that the initial orbit and the hyperbolic orbit share the same orbital plane).
2.2 Interplanetary Transfers
When in an interplanetary mission, the spacecraft is launched from a departure planet and accel- erated to a velocity higher than the local escape velocity such that it enters a hyperbolic arc to leave the vicinity of that planet. When the distance from the home planet is sufficiently large (in relation to its radius), the gravitational attraction from the Sun becomes the main force driving the motion of the probe, in its heliocentric trajectory at this point. This trajectory is designed according to the goals of the mission. While approaching the targeted planet, its gravitational attraction becomes more and more relevant until it eventually surpasses that of the Sun and becomes the main force acting on the spacecraft. At a small distance from the planet, the spacecraft can either just execute a flyby around it or be decelerated to enter an elliptical or circular orbit around it [7]. In a direct interplanetary trajectory the spacecraft leaves the departure planet and flies directly to the target one in a heliocentric transfer, not strongly interacting with any other celestial body in the Solar System (by entering its sphere of influence). On the other side, missions to the outer planets or even Mercury usually use non-direct trajectories where the probe executes one or more swing-bys (gravity
8 assists) along the way. For missions to Mercury these intend to decelerate the spacecraft so that is reaches closer distances to the Sun, while for missions to the outer planets these manoeuvres are employed to accelerate it to reach such great distances. The Sphere of Influence (SoI) of a planet is the region of space around that planet in which its gravi- tational attraction is greater than that of the Sun [14]. Laplace proven that such volume is approximately a sphere centred in the planet and with radius
2/5 mp RSoI ≈ ρ , (2.16) mSun where mp and mSun are, respectively, the masses of the planet and the Sun and ρ is the distance between them [7]. For the inner planets RSoI is less than 1% of the distance from the Sun to the planet, while for the outer planets such percentage increases up to 6.2% for Jupiter. Nevertheless, it can still be assumed as a first order approximation that in the heliocentric reference frame this radius is negligible, being contracted to a single point at the position of the planet. When comparing to the radius of the planets, RSoI is much larger, changing between 46 times for Mercury and 350 times for Neptune. If a complete and precise determination of the motion of the spacecraft is to be obtained, then a many-body problem must be solved using numerical methods due to its high complexity. Nevertheless, for preliminary designs of interplanetary missions such as those considered here, a simpler model is to be applied which not only produces solutions faster than using a more complex model but also provides better physical insight about the problem itself. This model is the patched conic approximation in which the motion of the spacecraft is divided into different segments/conics depending on whether the probe is within the sphere of influence of a planet (or other celestial body in the Solar System) or in the heliocentric phase of its trajectory (outside any sphere of influence) [7, 14]. Other simplifications can be added to obtain even simpler models, for example, it can be assumed that all planets orbit the Sun in the same plane, meaning that the problem is reduced from three to two dimensions. For an interplanetary transfer from Earth to a planet, the spacecraft’s trajectory is, according to the patched conic approximation, a succession of three two-body problems, being described as follows [7]:
1. Inside the Earth’s sphere of influence, a non-rotating planetocentric reference frame is used to study not only the launch and/or the parking orbit in which the probe is initially, but also the hyper- bolic orbit to leave the vicinity of the planet. At a great distance, where it crosses the sphere of
influence, the velocity of the spacecraft is almost equal to v∞.
2. Despite being very large when compared to the radius of the Earth, the sphere of influence in the heliocentric reference frame it is very small; so, it is assumed to be a single point. After crossing
the sphere of influence with velocity ~v ≈ ~v∞, the velocity of the spacecraft in this heliocentric reference frame is
~vE = V~E + ~v∞ (2.17)
where V~E is Earth’s velocity in that same frame. From this moment, the space probe travels in a
heliocentric trajectory until it reaches the sphere of influence of the target planet with velocity ~vp.
9 3. After crossing the sphere of influence of the planet, a new change in reference frame is performed to a non-rotating planetocentric one. Here, the spacecraft approaches the planet in a hyperbolic arc whose velocity at infinity can be defined as
~v∞ = ~vp − V~p , (2.18)
being V~p the velocity of that planet in the heliocentric frame. After reaching the periapsis of the hyperbolic orbit, the spacecraft can either be injected in a final elliptical/circular orbit around the planet or just perform a flyby, leaving the planet in that same hyperbolic trajectory.
As indicated by the name of the model, the different arcs (conics) flown by the spacecraft and which compose the overall interplanetary trajectory must be patched in order to obtain a somewhat continu- ous trajectory from the departure planet till the arrival at its final destination. This is done through an appropriate choice of orbital parameters and epochs for departure, possible encounters and arrival at the target planet. The key parameters in the design of an interplanetary mission are the target planet, the desired flight time and the launch and/or departure windows [14]. Moreover, for a mission departing from Earth the spacecraft can be directly injected into a hyperbolic orbit by the upper stages of the launcher or be inserted in a Low Earth Orbit (LEO) from which it after injects itself in the departing orbit. The design of the mission is a fundamental step in its performance since it defines the spacecraft mass and, consequently, its size [15]. Although a complete three-dimensional motion is considered in this thesis, by using the 3D ephemeris of the planets in the Solar System provided by JPL (Jet Propulsion Laboratory) [16], it is assumed that the initial/final orbit is in the same plane as the departing/approaching hyperbolic orbits defined by the heliocentric transfer phase.
2.3 Orbital Manoeuvres
Orbital manoeuvres can be distinguished in two main categories: change in the translational motion (orbit) of the spacecraft or change in its rotational motion to control the attitude of the spacecraft. Due to the nature of this work, only manoeuvres intended to change the orbits of the spacecraft are considered. In addition, it is assumed that all of them consist of impulsive manoeuvres, meaning that the change in the state of the spacecraft, more properly in its velocity vector, occurs instantaneously. This assumption is quite reasonable as a first order approximation when using high thrust propulsion systems where the thrust provided is considerably large. The thrust required to accelerate or decelerate the spacecraft can either be applied closer to a planet or another celestial body, i.e., inside its sphere of influence, or it can be applied when the spacecraft is in a heliocentric transfer. The latter case corresponds to the so called Deep Space Manoeuvres (DSM) and they are mainly used for corrections in the trajectory of the probe so that it correctly reaches its destination.
10 An impulse provided by the thrusters changes the velocity of the spacecraft in ∆~v, without changing its position, at the cost of consuming fuel and the mass of the spacecraft suddenly drops. If the mass immediately before the impulse is equal to m0 then, after the manoeuvre, the final mass of the spacecraft and the fuel mass consumed are given by Tsiolkovsky’s rocket equation [17]:
∆v mf = m0 exp − (2.19a) g0Isp ∆v ∆mfuel = m0 1 − exp − , (2.19b) g0Isp
2 where g0 = 9.80665 m/s is the gravitational acceleration near the surface of the Earth and Isp (in seconds) is the specific impulse of the thrusters.
For an impulsive manoeuvre, where the position remains the same during the instantaneous thrusting phase, the change in the orbital energy per unit of mass of the spacecraft, its specific energy, is equal to the change in its kinetic energy [7]:
1 1 ∆E = ∆K = (v2 − v2) = (∆v)2 + ~v · ∆~v, (2.20) 2 f i 2 i which means that the manoeuvre is more efficient, i.e, it requires less thrust for higher change in specific energy, if the impulsive shot is applied when the velocity of the spacecraft ~vi is maximum and it is directed along the direction of that velocity (the same direction to accelerate and opposite direction to brake). Therefore, these manoeuvres should be applied at the periapsis of the orbits for maximum efficiency.
If a change in orbital plane is required, which is a manoeuvre that usually requires a great amount of
∆v and propellant mass, then the expression for the impulsive shot required to change from ∆~vi to ∆~vf , which make an angle between them equal to γ, is given by [14]:
q 2 2 ∆v = vi + vf − 2vivf cos γ . (2.21)
It can be easily proved from Equation (2.21) that if the initial and final velocities have the same magni- tude, the expression for the impulse necessary simplifies to
γ ∆v = 2v sin . (2.22) i 2
Equation (2.22) shows that a pure change in the orbital plane, without changing the magnitude of the velocity, should be executed whenever the velocity is minimum since the impulse required is proportional to that velocity. Also, the bigger the angle γ, the costly the manoeuvre gets.
For a change in magnitude and direction, the optimum solution is very close to a complete change in magnitude at the periapsis where the velocities are maximum and a complete change in direction at the apoapsis for elliptical orbits or at an infinite distance for hyperbolic orbits. Theoretically speaking, a plane change at a infinite distance in a parabolic arc would not require any impulsive manoeuvre since vi = v∞ = 0.
11 2.4 Gravity Assist Manoeuvres (Swing-bys)
Planets and other celestial bodies in the Solar System, apart from the Sun, can be used to increase or decrease the heliocentric energy and velocity of a spacecraft due to an exchange of momentum between both bodies. The conservation of momentum dictates that for such gravitational interaction the following expression must remain valid [18]:
i ~ i f ~ f ms~vs + mpVp = ms~vs + mpVp , (2.23)
where ms and mp are, respectively, the masses of the spacecraft and of the planet and ~vs and V~p are their velocities in an ”inertial” reference frame (heliocentric). The index i refers to the moment immediately before the interaction while the index f refers to the moment immediately after. Equation (2.23) shows that not only does the heliocentric velocity of the spacecraft changes, but also that of the planet. Nevertheless, since mp ms, the change in V~p is residual. Three main categories of gravity assists exist: unpowered, powered and aerogravity assist [18]. In the first case, the only force acting on the spacecraft during the interaction is the gravitational pull from the planet. On the other side, a powered gravity assist manoeuvre also includes the application of thrust and an aerogravity assist uses the atmosphere of the planet to help changing the energy and velocity of the spacecraft. Only the first two cases will be further considered in this thesis.
2.4.1 Unpowered Gravity Assist
In the case of an unpowered swing-by, where no thrust is applied during the entire manoeuvre, the spacecraft flies a unique hyperbolic trajectory around the planet and energy conservation in the plan- in out etocentric reference frame implies that ||~v∞|| = ||~v∞ ||. Despite having the same magnitude, those velocities at an ”infinite” distance from the planet have different directions and it is that difference that leads to a change in the orbital energy of the probe in the heliocentric frame. This can be better under- stand by looking at the velocity vector diagram presented in Figure 2.1.
Figure 2.1: Velocity vector diagram for an unpowered gravity assist, taken from Melman [18]. The velocities ~v∞ correspond to the velocities of the spacecraft in the planetocentric reference frame upon crossing the SoI of the planet, with heliocentric velocity V~p. The vectorial sum of those velocities leads to the heliocentric velocities of the spacecraft ~vs immediately before and after the swing-by.
12 In the heliocentric reference frame, this manoeuvre is assumed as an ”instantaneous” change in the state of the spacecraft through a change in its velocity at the point where the encountered planet is. However, when zooming in to the planetocentric reference frame, a more complex geometry arises, as shown in Figure 2.2.
Figure 2.2: In-plane geometry for an unpowered swing-by as viewed in the planetocentric frame, taken from Melman [18]. The hyperbolic arc flown by the spacecraft inside the planet’s sphere of influence is presented in red. B~ represents the impact parameter, ~rp points to the periapsis of the orbit where the velocity of probe is ~vp. The angle α is the deflection of ~v∞ during the manoeuvre and β is the angle between the direction of motion of the planet around the Sun and the direction of the periapsis of the hyperbolic arc.
Using the patched conic approximation for the mission’s design, no information about where the interplanetary trajectory of the spacecraft intersects the sphere of influence of the planet can be obtained and, consequently, no information about the impact parameter B (closest distance of the planet to the incoming assymptote) or radius of periapsis rp can be directly obtained. So, these two parameters, related by the expression [7] s 2 4 ! µ B v∞ rp = 2 1 + 2 − 1 , (2.24) v∞ µ are usually left as optimization variables.
By choosing a value for the radius of periapsis rp (or the respective impact parameter B), the following expression holds for the turning (or deflection) angle α of the hyperbolic excess velocity ~v∞ during the manoeuvre [7] α 1 1 1 sin = = 2 = , (2.25) rpv∞ q B2v4 2 e 1 + ∞ µ 1 + µ2 where e is the eccentricity of the hyperbolic orbit. This means that the heavier the planet is or closer the swing-by occurs, the more ~v∞ will bend.
13 The change in the heliocentric specific energy of the spacecraft during the manoeuvre is given by [7]
2Vpv∞ cos β ∆E = 2 ; (2.26) rpv∞ 1 + µ using β = 0 and the periapsis radius equal to the radius of the planet Rp, the maximum theoretical change in specific energy is 2Vpv∞ ∆Emax = 2 . (2.27) Rpv∞ 1 + µ Figure 2.3 presents a plot with the maximum change in the specific orbital energy of a spacecraft after swing-by at five different planets in the Solar System as a function of ||~v∞||. Although it is not plotted, the curve for the Earth is very similar to that of Venus, being always just a little bellow. It can be noticed that the best planets for gravity assist manoeuvres in terms of energetic efficiency are Jupiter, Venus, Saturn and Earth itself.
Figure 2.3: Maximum change in the specific energy of a spacecraft during a swing-by, taken from Cor- nelisse et al. [17] and obtained using Equation (2.27). Only five planets are represented: Venus, Mars, Jupiter, Saturn and Uranus. The curve for Earth follows closely the curve of Venus, being always a little bellow it. Mercury and Neptune are not represented here since they are not very efficient due to the low mass of Mercury and the low orbital velocity of Neptune.
2.4.2 Powered Gravity Assist
If the incoming and outgoing hyperbolic excess velocities are perfectly match in terms of magnitude, in out ||~v∞|| = ||~v∞ ||, and the required turn angle αreq can be provided by just choosing the right radius of periapsis, without going bellow a certain minimum safety radius of periapsis r˜p, then an unpowered in out swing-by is sufficient to match the incoming and outgoing hyperbolic arcs. However, if ||~v∞|| 6= ||~v∞ || and/or the required turn angle is bigger than the maximum value αmax obtained from Equation (2.25) with rp =r ˜p, thrust is necessary to match both arcs, leading to a powered swing-by, whose geometry in the planetocentric frame is presented in Figure 2.4.
14 Figure 2.4: In-plane geometry for a powered swing-by as viewed in the planetocentric frame, taken from Melman [18]. The trajectory now consists of two different hyperbolic arcs matched at the periapis ~rp common to both of them. The half-turn angles from the arrival and departing orbits are now different.
Assuming for now that αreq ≤ αmax, it is only necessary to match the magnitude of the hyperbolic excess velocities of the spacecraft when crossing the sphere of influence of the planet. This match is in entirely made in the periapsis of the trajectory at an altitude such that α = αreq. Since α /2 is different out from α /2, a more complex expression relating α and rp is obtained starting from Equation (2.25):
αin αout 1 1 α = + = arcsin in 2 + arcsin out 2 . (2.28) 2 2 rp(v∞ ) rp(v∞ ) 1 + µ 1 + µ
This deflection α, as given by Equation (2.28), must be equal to the angle between those incoming and outgoing velocities , i.e., in out ~v∞ · ~v∞ α = arccos in out . (2.29) ||~v∞||.||~v∞ ||
To find α from Equations (2.28) and (2.29), an iterative process is required due to the presence of the arcsine functions. Once rp is found, being greater than r˜p because of the assumption of having
αreq ≤ αmax, the impulsive shot required at periapsis is easily obtained from the equations derived for hyperbolic orbits in Section 2.1.3:
s s 2µ 2µ out in out 2 in 2 ∆vp = |vp − vp | = + (v∞ ) − + (v∞) . (2.30) rp rp
If, unlike in the previous case, it is obtained a periapsis altitude smaller than the minimum safety value, because of having αreq greater than αmax obtained from Equation (2.28) with rp =r ˜p, a second impulsive shot must be provided and added to the already obtained ∆vp to get the delta-V budget of the manoeuvre. This second impulse is applied immediately after entering or immediately before leaving the sphere of influence of the planet, where the velocity is smaller, and it serves to rotate the incoming or the outgoing hyperbolic excess velocity by an angle δα = αreq − αmax, so that the swing-by can be correctly executed at the minimum altitude allowed. The magnitude of this second shot is given by
δα ∆v = 2vin/out sin , (2.31) δα ∞ 2
in/out where v∞ is the smallest of both hyperbolic excess velocities.
15 2.5 Lambert’s ”Orbital Boundary-value Problem”
Lambert’s ”orbital boundary-value problem”, also known as Lambert’s problem, is an important result in orbital mechanics and it is widely used in the design of space missions. In simple words, it consists of finding the orbit of a body under an inverse-square central gravitational force (when a point mass is considered) that brings it from point P1 to point P2 in a certain amount of time ∆t [19]. It becomes clear why it is considered a boundary-value problem: knowing the boundaries of the trajectory, it is desired to find the trajectory itself. A simple sketch of a solution corresponding to an elliptical orbit passing through points P1 and P2 is presented in Figure 2.5.
Figure 2.5: Example of an elliptical solution for Lambert’s problem with certain boundary conditions, taken from Battin and Vaughan [19]. The arc to be flown starts in P1 and ends in P2 and takes a certain amount of time ∆t. The main body is located in the focus F of the ellipse. Also, c is the chord connecting both positions and θ is the central angle.
For interplanetary missions, this result plays a fundamental role in the preliminary stages of the mission design, more properly in the heliocentric phases of the trajectory. If the position of the departure planet at t1 is ~r1 and the position of the target planet at t2 is ~r2, then the solution of this particular Lambert’s problem corresponds to the orbit that the spacecraft has to travel around the Sun so that it rendezvous with the target planet at epoch t2; therefore, the transfer time is ∆t = t2 − t1. The original theorem, which is just a reformulation of Kepler’s equation for elliptical orbits [7], was discovered by J.H. Lambert in 1761 and was only applicable to elliptical orbits. It can be stated as follows: the time required to travel an elliptical arc depends exclusively on the semi-major axis a of the orbit, the sum of the radius of both positions r1 + r2 and the distance between those positions c = ||~r2 − ~r1|| [7]. This theorem was only mathematically proven in 1778 by J.L. Lagrange. Since then, many have studied this problem with the intention of making it more general and with applications to the actual design of space missions, especially with the beginning of the space era in the middle of the 20th century [20]. Among those contributions, special emphasis should be given to
16 the work of Lancaster and Blanchard [21], presented in the late 60s, for its unified form of the problem including also a simple procedure for its resolution. Later, Gooding [22] built upon that work to present a high precision and more efficient procedure for all possible geometries of the problem, requiring only three iterations to obtain that high resolution. Finally, Izzo [20] published in 2014 an even better algorithm to solve the problem, built upon those two works mentioned, using a different technique for choosing the initial guesses of the parameters used in the iterative process, making it even more efficient.
2.5.1 Izzo’s Final Form of Lambert’s Problem
From all the different approaches existing to solve Lambert’s problem, either referred in the last paragraph or not, the one that has been chosen to use in this thesis is that of Izzo [20]. The rea- son for this selection is that it is very general, considering elliptical, parabolic and hyperbolic solutions and single-revolution and multi-revolution solutions, being also one of the most efficient in terms of the computational cost required to obtain high precision results. Moreover, this specific Lambert solver is already implemented and tested in PyKEP,simplifying the implementation of the code as presented later in Chapter5. Only the main results and equations for this final form of the problem are presented here. The reader is invited to read Izzo’s article, mentioned in the last paragraph, to get a complete derivation of the procedure if interested in that. In addition, the equations derived and presented here consider both elliptical and hyperbolic orbits. In the case of parabolic orbits, Battin’s work [19] is used by the author to obtain the solution and can be also seen in both articles, not being presented here. The derivation starts from Kepler’s equation for the transfer time between two points for both elliptical and hyperbolic orbits:
q a3 (E2 − E1 + e cos E1 − e cos E2 + 2Mπ) (elliptical) ∆t = t − t = µ (2.32) 2 1 q −a3 µ [e cosh H2 − e cosh H1 − (H2 − H1)] (hyperbolic) where M is the number of complete revolutions (M ≥ 0 for elliptical orbits and M = 0 for parabolic and hyperbolic ones) and Ei and Hi are, respectively, the eccentric and hyperbolic anomalies at epoch ti and related to the true anomaly θi by [7]
E r1 − e θ tan i = tan i (2.33a) 2 1 + e 2 H re − 1 θ tanh i = tanh i . (2.33b) 2 e + 1 2
The final equation to be solved makes direct and indirect use of several parameters related to the boundary conditions of the problem. One of them is the semi-perimeter of the triangle formed by the position vectors ~r1, ~r2 and the vector ~c connecting them:
r + r + c s = 1 2 . (2.34) 2
17 A non-dimensional parameter λ is also used and it is related to that semi-perimeter s and chord c of the triangle by √ r r θ r c λ = 1 2 cos = 1 − ; (2.35) s 2 s
λ ∈ [−1, 1] is positive for θr ∈ (0, π) and negative for θr ∈ (π, 2π), where θr is the transfer angle reduced to the interval (0, 2π).
The transfer time ∆t is also made non-dimensional by using
r2µ T = ∆t (2.36) s3 in order to obtain a completely non-dimensional equation for the final form of the problem. Also, the different parameters s, λ and T can be known a priori from the conditions of the problem.
With all those parameters defined, the final form of the problem is given by the following equation:
! 1 ψ + Mπ T = − x + λy , (2.37) 1 − x2 p|1 − x2| where x is the variable to which the problem is solved, being also non-dimensional and related to the semi-major axis of the orbital solution by s x = 1 − . (2.38) 2a
This variable lies in the interval (−1, 1) for elliptical orbits and is bigger than 1 for hyperbolic ones. For x = 1 the solution is a parabolic arc but a different method should be used due to the cancellation of the numerator and denominator in Equation (2.37), leading to a lack of precision in the results.
The non-dimensional auxiliary variables y and ψ, used to simplify the writing of Equation (2.37), are obtained from the variable x and the parameter λ. The former is simply given by y = p1 − λ2(1 − x2), while the latter must be computed regarding the type of orbit: for elliptical solutions it is obtained from
cos ψ = xy + λ(1 − x2) (2.39a) p sin ψ = (y − λx) 1 − x2 (2.39b) and for hyperbolic solutions it is obtained from
cosh ψ = xy − λ(x2 − 1) (2.40a) p sinh ψ = (y − λx) x2 − 1 . (2.40b)
Once the solution x, and corresponding value for y, is obtained through an iterative process, as explained later in Section 2.5.3, the actual orbit to be flown can be computed by knowing the initial and
final positions (~r1 and ~r2) from the problem’s definition and the initial and final velocities (~v1 and ~v2) computed from
~vi = vr,i~ir,i + vt,i~it,i , (2.41)
18 where the unit vectors are obtained from ~r1 and ~r2, as explained in Izzo [20]. Besides those unit vectors, the other parameters in Equation (2.41) are obtained using the following expressions:
vr,1 = γ[(λy − x) − ρ(λy + x)]/r1 (2.42a)
vr,2 = −γ[(λy − x) + ρ(λy + x)]/r2 (2.42b)
vt,1 = γσ(y + λx)/r1 (2.42c)
vt,2 = γσ(y + λx)/r2 , (2.42d)
p p 2 where γ = µs/2, ρ = (r1 − r2)/c and σ = 1 − ρ are just auxiliary parameters.
2.5.2 Single-revolution versus Multi-revolutions
To help understand the solution space of Lambert’s problem as derived by Izzo in terms of the universal variable x, it is presented in Figure 2.6 the curves for T with respect to x for different values of the parameter λ and different values of complete revolutions (M ≥ 0). It is clearly noticed how multi-revolutions can only occur for elliptical solutions, where |x| < 1. A unique solution is always obtainable for the single-revolution case (M = 0) with λ 6= 1. However, in a multi-revolution case, with a certain value for M, there are zero, one or two solutions depending on M T and Tmin, which is the minimum transfer time needed for the existence of a solution with M complete revolutions and whose value increases with M. If T < Tmin, no solution exists. On the other hand, if T > Tmin then two solutions x1 and x2 (with x1 < x2) are obtainable. In the limit case of having
T = Tmin, both solutions converge to xTmin , leading to a single solution for the problem. It is easily shown that if Mmax is the maximum value for which solutions exist, then the total number of solutions for
Mmax Mmax that particular problem is 2Mmax + 1 if T > Tmin or 2Mmax if T = Tmin .
Figure 2.6: Non-dimensional transfer time curves as a function of x for different values of λ and M, taken from Izzo [20]. The different curves correspond to the solution of Equation (2.37) for different values of the parameters λ and M and of the variable x.
19 2.5.3 Izzo’s Approach to Solve Lambert’s Problem
Equation (2.37) cannot be solved implicitly and, for that reason, an iterative process must be adopted to obtain an approximate solution x˜ as close as possible to the actual solution x. Using x as the iterative variable, Householder iterations are employed to obtain the solution of the problem, rewritten now as
! 1 ψ + Mπ f(x) = − x + λy − T = 0 , (2.43) 1 − x2 p|1 − x2| where y and ψ are found from x during each iteration using the equations presented in Section 2.5.1. Starting from an initial guess x0, at each iteration the new value of x is obtained from the previous one using [f 0(xk)]2 − f(xk)f 00(xk)/2 xk+1 = xk − f(xk) , (2.44) [f 0(xk)]3 − f(xk)f 0(xk)f 00(xk) + f 000(xk)[f(xk)]2/6 where the first three derivatives of f with respect to x are computed recursively as follows:
df 1 x f 0 = = 3x(f + T ) − 2 + 2λ3 (2.45a) dx 1 − x2 y d2f 1 1 f 00 = = 3(f + T ) + 5xf 0 + 2λ3(1 − λ2) (2.45b) dx2 1 − x2 y3 d3f 1 x f 000 = = 7xf 00 + 8f 0 − 6λ5(1 − λ2) . (2.45c) dx3 1 − x2 y5
Knowing the iteration variable and the iterative algorithm, the only step left is the definition of the initial guess for x, which depends on considering a single-revolution or a multi-revolution case. This is the most relevant improvement that Izzo brought, leading to a more efficient algorithm. For the single-revolution case, where only one solution exists, the initial guess for the iteration variable should be determined by 2/3 T0 T − 1 if T ≥ T0 0 x = 5T1(T1−T ) , 5 + 1 if T ≤ T1 (2.46) 2T (1−λ ) log (T /T ) T0 2 1 0 T − 1 if T1 < T < T0 √ 2 being T0 = arccos λ + λ 1 − λ the value of T obtained from Equation (2.37) with x = 0 (and M = 0) 3 and T1 = 2(1 − λ )/2 the value of T for x = 1 obtained from Battin and Vaughan [19]. Contrary to the previous situation, the multi-revolution case brings two solutions to the same bound- ary conditions and for a specific number of complete revolutions M, assuring that the non-dimensional transfer time T is bigger than a certain minimum value Tmin, for which a single solution exists and bel- low that no solutions can be obtained. Assuming that T > Tmin, i.e., two solutions exist, then the initial guesses to search for the first and second solutions (with x1 < x2) are obtained from