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Research Papers

Hybrid Electric Vehicle Fault Tolerant Control OPEN ACCESS

[+] Author and Article Information
Richard T. Meyer

Department of Mechanical and
Aerospace Engineering,
Western Michigan University,
Kalamazoo, MI 49008
e-mail: richard.meyer@wmich.edu

Scott C. Johnson

School of Electrical and
Computer Engineering,
Purdue University,
West Lafayette, IN 47907
e-mail: johns924@purdue.edu

Raymond A. DeCarlo

School of Electrical and
Computer Engineering,
Purdue University,
West Lafayette, IN 47907
e-mail: decarlo@ecn.purdue.edu

Steve Pekarek

School of Electrical and
Computer Engineering,
Purdue University,
West Lafayette, IN 47907
e-mail: spekarek@purdue.edu

Scott D. Sudhoff

School of Electrical and
Computer Engineering,
Purdue University,
West Lafayette, IN 47907
e-mail: sudhoff@purdue.edu

1Corresponding author.

Contributed by the Dynamic Systems Division of ASME for publication in the JOURNAL OF DYNAMIC SYSTEMS, MEASUREMENT, AND CONTROL. Manuscript received September 26, 2016; final manuscript received June 13, 2017; published online September 8, 2017. Assoc. Editor: Azim Eskandarian.

J. Dyn. Sys., Meas., Control 140(2), 021002 (Sep 08, 2017) (12 pages) Paper No: DS-16-1461; doi: 10.1115/1.4037270 History: Received September 26, 2016; Revised June 13, 2017

This paper investigates the supervisory-level, fault tolerant control of a 2004 Prius powertrain. The fault considered is an interturn short circuit (ITSC) fault in the traction drive (a surface mount permanent magnet synchronous machine (SPMSM) for which its rotor is part of the vehicle's driveline). ITSC faults arise from electrical insulation failures in the stator windings where part of a phase winding remains functional while the remaining decoupled windings form a self-contained loop. Because the permanent magnets on the rotor (driveline) shaft are able to induce very large eddy currents in this self-contained loop if its rotational velocity is left unchecked, the maximum allowable driveline speed, and consequently, vehicle speed, must be reduced to avoid exceeding the drive's operational thermal limits. A method for detecting these ITSC faults and the induced eddy current in an SPMSM using a moving horizon observer (MHO) is reviewed. These parameters then determine which previously computed, fault-level-dependent SPMSM input–output power efficiency map and maximum safe operating speed is utilized by the supervisory-level controller. The fault tolerant control is demonstrated by simulating a Prius over a 40 s drive velocity profile with fault levels of 0.5%, 1%, 2%, and 5% detected at the midpoint of the profile. For comparison, the Prius is also simulated without a traction motor fault. Results show that the control reduces vehicle velocity upon detection of a fault to an appropriate safe value.

FIGURES IN THIS ARTICLE
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Hybrid electric vehicles, such as the Toyota Prius, use both an internal combustion engine (ICE) and electric machine, e.g., permanent magnet synchronous machine (PMSM), for propulsion. When a fault occurs in a PMSM, it commonly is an interturn short circuit (ITSC) in the stator windings [1]. Continued electrical excitation of the PMSM during a fault results in hot spots that may lead to fire [2,3]. Electrical excitation can result from the PMSM inverter as well as rotor spin due to vehicle movement. In the latter case, stopping the vehicle to eliminate excitation may leave the driver far from service or in an unsafe location.

A review of pertinent PMSM fault detection work is in Refs. [2,3]. With regards to PMSM fault mitigation after fault detection, actions are dividable into hardware reconfiguration and fault tolerant control [4]. Hardware reconfiguration usually requires some redundancy or additional component features that can be activated during a fault to reduce/bypass its effect. This approach requires additional hardware cost [5,6]. Fault tolerant control can be classified as either passive or active [5]. In the passive case, the control is designed to be robust to certain faults. In contrast, active control, initiated by a fault detection capability, modifies component operation to mitigate the effects of the fault. Past work in active fault tolerant control is primarily focused on using observers to replace information from faulty motor sensors [710]. However, modification of hybrid electric vehicle control strategies during a motor fault is considered in Phillips et al. [11] and Xu et al. [12]. In both works, vehicle operation is restricted in some way with respect to the fault, as for example, by reducing maximum allowable motor power. Neither of these works specify how variations in the level of motor fault are incorporated into the vehicle powertrain control.

Herein, a moving horizon observer is used to identify the presence of a fault and the fault level in a 2004 Toyota Prius traction PMSM. Given this fault information, the next objective is to develop fault mitigation controller strategies that allow PMSM-based hybrid vehicles to continue operation albeit at a substantially reduced level. Fault mitigation is carried out at the so-called supervisory level. A supervisory-level controller coordinates vehicle control by determining optimized power flows to the individual subsystems such as was explored in Refs. [1315]. For example, for a diverse set of situations, the supervisory-level controller would determine how best to utilize the electric drive system in coordination with the ICE to meet driving objectives.

For efficient and feasible numerical optimization strategies, supervisory-level subsystem models are power flow based; specifically, the subsystem models are low granularity power flow component models that utilize efficiency maps as opposed to high-granularity models based on the underlying physics. In the case of the PMSM, such an efficiency map depends on whether or not the motor has a fault as well as on the degree of fault. An advantage of the “online” supervisory-level control is that it can adapt the controls to different fault levels as they occur without requiring an exhaustive library of precomputed controls.

This paper investigates: (i) the hybrid, i.e., switched system, supervisory control problem in the Toyota Prius; (ii) construction of efficiency maps and constraints for a PMSM with ITSC faults; (iii) work for PMSM ITSC fault detection using a moving horizon observer; and (iv) a high-level vehicle power flow control optimization strategy during both normal and ITSC fault operation.

Figure 1 shows the architecture of the 2004 Toyota Prius's hybrid powertrain with a 57 kW Atkinson-cycle ICE, a 21 kW nickel metal hydride (NiMH) battery pack, two SPMSMs,2 and a power split device (PSD) consisting of a planetary gear system that connects the power flow pathways of the ICE and SPMSMs. One SPMSM is for power generation and engine starting, labeled “SPMSM1,” with a maximum mechanical power of 30 kW and the other SPMSM is for traction and regenerative braking, labeled “SPMSM2,” with a maximum mechanical power of 50 kW.

Due to differing efficiency maps for SPMSM1 and SPMSM2 propelling and generating, the Prius architecture is modeled as a switched system with a finite set of “modes of operation.” Each mode represents a fixed configuration of power flow directions and is characterized by potentially different efficiency maps and dynamics: (i) SPMSM1 propelling (ICE start-up)-SPMSM2 propelling-battery discharging-ICE off; (ii) SPMSM1 generating-SPMSM2 propelling-battery discharging-ICE on; (iii) SPMSM1 generating-SPMSM2 propelling-battery charging-ICE on; and (iv) SPMSM1 generating-SPMSM2 generating-battery charging-ICE on or off.

In the remainder of this paper, we will consider an ITSC fault only occurring in SPMSM2, i.e., no faults occur in SPMSM1. SPMSM1, and SPMS2 models are developed in Sec. 4 and SPMSM2 fault detection is presented in Sec. 5. Section 6 describes the simulation results of the control and interturn short circuit fault. Section 7 concludes the paper. In the Appendix, the supervisory-level models are constructed for the ICE, battery, vehicle, power split device, and electrical bus. The Appendix also describes the interconnection structure and other modeling issues associated with the Prius.

As mentioned, a supervisory-level controller specifies power flows in and out of the subsystems/components. The low-granularity power flow control oriented models (as opposed to first principles high-granularity models) allow for feasible and more efficient numerical optimizations.

To notationally identify the modes set forth in Sec. 2, let αi ∈ {0, 1}, i = 1, 2, 3, 4 denote the active (αi = 1) or nonactive (αi = 0) state of each mode. Since only one mode can be physically active at a time, we impose the constraint α1 + α2 + α3 + α4 = 1. We note that values of αi vary over time and can change instantaneously; a minimum dwell time is assumed to avoid the possibility of infinitely fast mode switching.

Given the modes of operation, the supervisory-level optimization problem has the form Display Formula

(1a)mini=1,,4u,αi{0,1}J(x0,xref,tp,0,tp,f,x,u,α¯)
Display Formula
(1b)subjecttox(tp,0)=x0
Display Formula
(1c)1=α1+α2+α3+α4
Display Formula
(1d)x˙=i=14αifi(t,x,ui)
Display Formula
(1e)0=i=14αihi(t,x,ui)
Display Formula
(1f)0i=14αigi(t,x,ui)

In the problem statement:

  • Jα(x0,xref,tp,0,tp,f,x,u,α¯) is a convex performance index, wherein x0 is the state at the current time tp,0, tp,f is the final prediction horizon time, tp,f − tp,0 is the prediction horizon length, x(·)n is the supervisory-level state consisting of all pertinent dynamic subsystem variables, xref(t) is a reference state trajectory to be tracked over [tp,0,tp,f],u(·)m×4 is the composite of the mode-specific continuous control input vectors, and α¯=[α1,α2,α3,α4] is the aggregate of mode values such that at each instant of time the entries of α¯ sum to one and there is at most one nonzero entry.

  • x=[Pfuel,Pice,Ωice,W¯bat,ϒv] is the state vector, where Pfuel is the fuel power delivered to the engine, Pice is the ICE output power, Ωice is the square of the ICE angular speed, W¯bat is the normalized battery state of charge (SOC), and ϒv is an energy function equal to the square of the vehicle velocity. The speed variables are squared to represent energy, which is ordinarily proportional to the square of speed. Further details can be found in the Appendix.

  • ui=[ufueli,u1i,u2i,ubrki], wherein ufueli is the commanded normalized fuel in the ith mode of operation; u1i is the commanded fraction of maximum speed-dependent power for SPMSM1 in the mode of operation, which is electrical in mode 1 and mechanical in the remaining modes; u2i is the commanded fraction of maximum speed-dependent power for SPMSM2 in the mode of operation, which is electrical in modes 1, 2, 3, and mechanical in 4; and ubrki is the commanded fraction of maximum speed-dependent frictional braking power in the ith mode.

  • The performance index (PI), J(⋅), is one of two expressions depending on whether the engine is on or off (note that the engine on or engine off is a fuel dependent control action and does not represent two distinct modes of operation) Display Formula

    J(·)={Joff(·),ICEoff(2a)Jon(·),ICEon(2b)
    Joff(·)=tp,0tp,fCv(ϒvϒv,ref)2+i=14Cbrkiαi(Pbrki)2+Cu1i=14αi(u1i)2+Cu2i=14αi(u2i)2dt
    Jon(·)=Joff(·)+tp,0tp,fCSOC(W¯batW¯bat,ref)2+Cfuel(Pfuel)2dt
    where ϒv,ref is the squared reference velocity or energy to track, W¯bat,ref is the desired battery state of charge, Cv is the penalty weight on the vehicle energy error (and implicitly vehicle tracking error), Cbrki is the mode-dependent penalty weight on frictional braking use to promote battery regenerative braking, Cu1 and Cu2 are the penalty weights on SPMSM1 and SPMSM2 usage to encourage bang-bang solutions in the numerical optimization described shortly,3CSOC is the penalty weight on deviation of the battery state of charge from a desired value, and Cfuel is the penalty weight on ICE engine fuel consumption.

  • x˙=44αifi(t,x,ui) represents a composite of the mode specific subsystem dynamics without interconnections. Specifically Display Formula

    (3)[P˙fuelP˙iceΩ˙iceW¯˙batϒ˙v]=i=14αi[ffueli(Pfuel,Ωice,ufueli)ficei(Pfuel,Pice,Ωice)fΩicei(Pice,Pice,psdi)fbati(W¯bat,Pbati)fvi(Pbrki,Pwhl,psdi,P2,Li,ϒv)]
    wherein Pice,psdi is the ICE power routed into the power split device in the ith mode, Pbati is the ith mode battery discharge or charge power, Pbrki is the ith mode frictional braking power, Pwhl,psdi is the ith mode power delivered to the drive wheels from the power split split device, P2,Li is the ith mode SPMSM2 power to the drive wheels, and ffueli(·),ficei(·),fΩicei(·),fbati(·), and fvi(·) are the right-hand sides of Eqs. (A1), (A2), and (A6)(A8), respectively, in the Appendix. Also, for reference yi=[Pbati,Pbrki,P1,ei,P1,Li,P2,ei,P2,Li,Pice,psdi,Pwhl,psdi],i=1,2,3,4 is the vector of mode-specific algebraic variables that depend upon states and continuous control inputs.

  • 0 = hi(t, x, ui) describes the ith mode interconnections of the subsystem models and any other equality constraints pertinent to the optimization problem. Specifically, the constraints are described by Eqs. (A28)(A35) in the Appendix.

  • 0 ≥ gi(t, x, ui) represents the ith mode inequality constraints pertinent to the optimization problem, for example, bounds on states, inputs, algebraic variables, etc.

Comments on the Numerical Solution of the Optimization Problem.

The minimization problem in Eq. (1) is numerically ill-conditioned, since the optimization requires searching over all combinations of switching sequences. As discussed in Ref. [18], this is an NP-hard problem that exponentially increases with the number of modes. To avoid this complexity, the switched optimal control problem in Eq. (1) is replaced with the embedded optimal control problem (a partial relaxation of only the modes of operation), which has the form [18,19] Display Formula

(4a)mini=1,,4u,αi[0,1]J(x0,xref,tp,0,tp,f,xe,u,α¯)
Display Formula
(4b)subjecttox(t0)=x0
Display Formula
(4c)1=α1+α2+α3+α4
Display Formula
(4d)x˙e=i=14αifi(t,xe,ui)
Display Formula
(4e)0=hi(t,xe,ui),i=1,2,3,4
Display Formula
(4f)0gi(t,xe,ui),i=1,2,3,4

Observe the embedded problem state is denoted as xe to distinguish it from the original problem state and the algebraic constraints, Eqs. (4e) and (4f), are now satisfied in every mode. The distinguishing aspect of the embedded optimization problem relative to the original optimization problem is the latter's requirement αi ∈ {0, 1} is relaxed to αi ∈ [0, 1]. This relaxation transforms the original problem with combinatorial complexity into a traditional nonlinear optimization. When αi = 1 for some i, then the embedded problem solution in Eq. (4) reduces to the switched problem solution in Eq. (1). When αi ∈ (0, 1), one must project the modal solution onto the nearest physically realizable mode of operation using techniques described in Refs. [1719]. However, the basic idea is to project the embedded fractional value of αi onto the nearest integer value, while maintaining the constraint Eq. (4c).

It was proven in Ref. [19] that the switched state trajectories of x of Eq. (1) are dense in the set of embedded system trajectories xe in Eq. (4), so a given embedded trajectory can be approximated by a switched system trajectory to any degree of desired accuracy. Thus, Eq. (4) provides a numerically viable nonlinear optimization problem whose solution can be used to generate approximate solutions to Eq. (1). For additional details concerning the embedded optimal control problem, we refer the reader to Ref. [19].

The specific numerical solution uses fmincon within matlab, which requires the optimization problem to be discretized. Briefly, the performance metric is discretized using trapezoidal numerical integration and the dynamics are discretized using collocation. The reader is referred to earlier works for details [17].

Because there has been no supervisory-level, efficiency-based modeling development of the SPMSM that accounts for ITSC faults in the literature, its development is presented here. The remainder of the models are presented in the Appendix.

SPMSM1.

At the supervisory level, SPMSM1 is assumed to have fault-free operation with input-to-output efficiency-based representation Display Formula

(5)P1,L=ηSPMSM,prop,1(ω1,m,P1,inv)P1,inv

for propelling and for generating Display Formula

(6)P1,inv=ηSPMSM,gen,1(ω1,m,P1,L)P1,L

where P1,L is the mechanical power, P1,inv is the inverter power, η denotes efficiencies, which depend on the input power, and the rotational speed of the rotor, ω1,m, which depends on the SPMSM1 parameters and powertrain operation as set forth in the Appendix.

The efficiency maps are generated from a detailed model of the SPMSM1 operation given in Refs. [2,3], and [20]. Specifically Display Formula

(7)ηSPMSM1,prop(ω1,m,sc,P1,inv,sc)a1+b1ω1,m,sc+c1ω1,m,c2+d1P1,inv,sc+e1w1,m,scP1,inv,sc
Display Formula
(8)ηSPMSM1,gen(ω1,m,sc,P1,L,sc)a2+b2ω1,m,sc+c2ω1,m,sc2+d2P1,L,sc+e2w1,m,scP1,L,sc

where ω1,m,sc=ω1,m/(1000π/3),P1,inv,sc=P1,inv/3×104,P1,L,sc=P1,L/3×104, and the coefficients a1,…, e2 are set forth in the Appendix.

SPMSM2.

For SPMSM2 the efficiency terms include potential faulted operation depending on the fault level σ as follows: Display Formula

(9)P2,L=ηSPMSM2,prop(ω2,m,P2,inv,σ)P2,inv

for propelling and for generating Display Formula

(10)P2,inv=ηSPMSM2,gen(ω2,m,P2,L,σ)P2,L

Here, P2,L is the mechanical power, P2,inv is the inverter power, and ω2,m is the rotor rotational speed, which depends upon the vehicle speed. The fault-level-dependent efficiencies have a form similar to Eqs. (7) and (8)Display Formula

(11)ηSPMSM2,prop(ω2,m,sc,P2,inv,sc,σ)a1(σ)+b1(σ)ω2,m,sc+c1(σ)ω2,m,c2+d1(σ)P2,inv,sc+e1(σ)w2,m,scP2,inv,sc
Display Formula
(12)ηSPMSM2,gen(ω2,m,sc,P2,L,sc,σ)a2(σ)+b2(σ)ω2,m,sc+c2(σ)ω2,m,sc2+d2(σ)P2,L,sc+e2(σ)w2,m,scPL,sc

where ω2,m,sc=ω2,m/(200π),P2,inv,sc=P2,inv/5×104,P2,L,sc=P2,L/5×104. SPMSM2 parameters and efficiency fit coefficients a1,…, e2 are given in the Appendix.

Any ITSC fault results in an eddy current, ifs, induced in the faulted coils by the permanent magnets on the spinning rotor shaft. The magnitude of ifs depends on the fault level, σ. The ifs2R losses can cause overheating that exceeds the maximum thermal limits for safe operation. Thus, the specific fault level, σ, constrains the maximum safe vehicle speed. This limit then induces a constraint on the supervisory-level control strategy as evidenced through the safe operating regions of Fig. 2. The fault level of SPMSM2 as well as the induced eddy current, ifs, must be estimated from a dynamic observer, which is implemented in this work as a moving horizon observer (MHO) [21]. The estimated σ is used to determine which efficiency map is now applied at the supervisory level.

As mentioned, an MHO is used to estimate the fault level σ and current ifs. We assume for convenience a single fault in phase-a.4 To do this effectively, one needs to explicate the underling mathematical model of SPMSM2 as a function of the fault level. The model set forth does not include saturation effects that can be incorporated as an additional mode of operation; this inclusion is beyond the scope of this paper.

SPMSM2 Component Level Model.

The SPMSM2 stator voltage equations with a single ITSC fault in phase-a (modified from those in Ref. [16] for physical realizability of the mutual inductances) are Display Formula

(13)[vasvbsvcs0]=[(1σ)Rs0000Rs0000Rs0000σRs][iasibsicsifs]+Lf(σ)ddt[iasibsicsifs]+[eaebecef]

and Display Formula

(14)Lf(σ)=[(1σ)2L(1σ)M(1σ)Mσ(1σ)L(1σ)MLMσM(1σ)MMLσMσ(1σ)LσMσMσ2L]

where vζs and iζs denote the stator voltage and current in phase ζ = a, b, c. The back emf terms are given by Display Formula

(15)[eaebecef]=ω2,mλm[(1σ)cos(θr)cos(θr2π/3)cos(θr+2π/3)σcos(θr)]

where θr = θ2,mnp/2 is the electrical angular position of the rotor and np is the number of motor poles. Note that the fault loop back emf, ef, has the same phase angle as the back emf in the faulted phase-a. Finally, by Kirchoff's current law Display Formula

(16)ias+ibs+ics=0

It is assumed that iζs are sinusoidal with the form Display Formula

(17)ias=Iqacos(θr)+Idasin(θr)
Display Formula
(18)ibs=Iqbcos(θr2π3)+Idbsin(θr2π3)
Display Formula
(19)ics=Iqccos(θr+2π3)+Idcsin(θr+2π3)

where I and I are the current amplitudes determined by the inverter. Controlling currents of the form in Eqs. (17)(19) has the advantage that the derivatives of the currents are known in Eq. (13), making the computation of vζs algebraic. This observation is exploited in the development of the MHO in Sec. 5.2.

SPMSM2 Fault Moving Horizon Observer.

The MHO is the minimization of the following quadratic performance index5 over the time interval [tk − T,tk]: Display Formula

(20)J=tkTtk[(yMŷ)TQ(yMŷ)+wpv̂fs2]dt

where Q is a weighting matrix and v̂fs6 is the estimated fault loop voltage that must be driven to (approximately) zero during the estimation process; yM is a vector of measured voltages and currents and ŷ is the estimates of those voltages and currents. Specifically Display Formula

(21)yM=[vabM,vbcM,vcaM,iasM,ibsM,icsM]T
Display Formula
(22)ŷ=[v̂asv̂bs,v̂bsv̂cs,v̂csv̂as,îas,îbs,îcs]T

Henceforth, the superscript M indicates a measured variable and hatted variables are estimates. All the terms in ŷ depend upon the fault level, sinusoids in Eqs. (17)(19), the time derivatives of Eqs. (13) and (17)(19). Hence, the minimization of Eq. (20) can be taken over the fault level σ̂ and the associated piecewise constant magnitudes of the sinusoidal currents—Îqa,Îda,Îqb,Îdb,Îqc, and Îdc—subject to Eqs. (13)(19), (21), and (22).

Observer Implementation.

The minimization of the performance index, Eq. (20), is over a 50 ms horizon with two partitions, each of 25 ms. Successive minimizations have an overlap of one partition. The forward most partition of the current minimization becomes the trailing partition in the next optimization. Optimal values from the forward most partition of the current minimization are used to initialize both partitions in the next optimization.

The steps in the optimization over each horizon are: (i) obtain the measured values at each partition's boundaries over the horizon; (ii) generate initial estimates of I, I, and the fault level σ; (iii) compute values of the estimated currents using Eqs. (17)(19) and the their derivatives evaluated at the appropriate time points; (iv) evaluate Eqs. (14)(15); (v) determine the estimates of line-to-neutral voltages vζs from Eq. (13); (vi) set v̂fs equal to the last row of the right-hand side of Eq. (13); (vii) form ŷ using Eq. (22); and (viii) use a discrete-time approximation of the integral in the performance index to compute the cost. Steps (iii)–(viii) are performed and repeated until a minimal cost is achieved within a nonlinear optimization solver, such as fmincon within matlab. To continue, the horizon is incremented forward in time and the optimization steps are repeated.

For this application, the MHO observer correctly identified the fault which begins at the 20 s mark in the simulations detailed in Sec. 6. Fault detection was achieved within 50 ms. Subsequent fault and fault current estimates are then used to switch between precomputed efficiency maps and maximum safe operating speed limits, which modifies the supervisory-level power management optimization problem. Results follow in Sec. 6.

This section reviews the performance of the Prius over a 40 s trapezoidal drive profile subject to an ITSC fault at 20 s. The trapezoidal profile consists of: (i) a constant acceleration increase in velocity from rest to 20.12 m/s (45 mph) over [0,10)s, (ii) constant velocity over [10, 30)s, and (iii) constant deceleration to rest over [30, 40]s. Figure 3 shows the commanded trapezoidal and the simulated unfaulted vehicle response on a level road, i.e., θv = 0.

Drive profile control simulations are performed using techniques set forth in Ref. [17]. The optimization problem is discretized with a Tc = 0.25 s interval time partition according to the methods in Ref. [17]. In the PI, Eq. (2a), the prediction horizon is 0.5 s (two partitions) and the penalty weights are Cv = 100, Cbrki=1000/(50)2 for i = 1, 2, 3, Cbrk4=100/(50)2,Cu1=2,Cu2=2,Cfuel=10/(383.5)2, and CSOC = 1000, when W¯bat<W¯bat,ref=0.58 and zero otherwise.

We note that the vehicle cannot change velocity instantaneously, and thus, a velocity reference obtained from the operator at kTc becomes the commanded velocity to achieve at (k + 1)Tc, i.e., vref((k + 1)Tc). In the optimization, at time kTc, the present, the desired velocity at (k + 1)Tc is known. The value at (k + 2)Tc is unknown and is estimated linearly as Display Formula

(23)vref,est((k+2)Tc)=vref((k+1)Tc)+Tc[vref((k+1)Tc)vref(kTc)Tc]

Thus Display Formula

(24)ϒv,ref(kTc)=[vref(kTc)]2
Display Formula
(25)ϒv,ref((k+1)Tc)=[vref((k+1)Tc)]2
Display Formula
(26)ϒv,ref,est((k+2)Tc)=[vref,est((k+2)Tc)]2

The reference velocity is extrapolated from known current and past values that leads to the extrapolated reference, ϒv,ref,est, in Eq. (26). We use extrapolated values because we do not assume that future values of the drive profile are known. This linear extrapolation assumption is meant to approximate a driver but does add a small error to the tracking of reference signals that are nonpiecewise linear or have “corners.”

The Prius model is simulated again over the trapezoidal drive profile now with ITSC faults of 0.5% (σ = 0.005), 1% (σ = 0.01), 2% (σ = 0.02), and 5% (σ = 0.05) from 20 s onward. Figure 4 shows the vehicle velocity tracking achieved for each fault level. Every fault is considered to have been detected by the observer of Sec. 5 at t = 20 s. At 20 s, ϒv,ref is chosen to decelerate the vehicle at −2 m/s2 until ϒv,ITSC,max (the square of 95% of the maximum fault operation velocity) is reached. Then, the reference is the minimum of ϒv,ref and ϒv,ITSC,max until the fault is removed. During the deceleration phase, SPMSM2 is unpowered to avoid exacerbating thermal runaway and hence unsafe operating conditions. This velocity modification behavior during faults is exhibited in Fig. 4 for all of the fault cases.

Figure 5 shows the mechanical power (positive if propelling and negative if generating) of SPMSM1 and SPMSM2 over the drive profile with the 5% ITSC fault. Similar SPMSM1 and SPMSM2 responses are displayed with the other fault levels. SPMSM1 is off until it is used to start the engine from 3 to 3.5 s. Thereafter, SPMSM1 is driven by the ICE and sends power to the electrical bus to drive SPMSM2 and recharge the battery. After the fault occurs, SPMSM1 is on for the next 2 s, while the vehicle speed decreases and then remains off for the remainder of the drive profile. We note from 20.25 s onward that the ICE is off as well. SPMSM1 is still able to provide battery charge power from 20.25 to 22 s, during which the ICE is off, by consuming the ICE inertia energy. SPMSM2 propels the vehicle during the initial commanded acceleration from 0 to 10 s and both propel and recharge the battery from 10 to 20 s during the commanded constant velocity. Upon the detection of the fault, SPMSM2 is off, while the vehicle speed is decreased to the maximum safe fault operation velocity. Next, the drive is used to propel the vehicle at constant velocity from 26.5 s until the commanded deceleration to rest is encountered at 38 s, and then the drive provides battery regenerative braking to 39.75 s.

For reference, the ICE output power for the unfaulted and 5% fault-level cases are provided in Fig. 6. In addition, the projected mode selections are given in Fig. 7.

In this paper, we have reviewed a moving horizon observer to detect ITSC faults in a SPMSM, set forth the process to develop efficiency maps for an SPMSM with and without an ITSC fault and demonstrated how the knowledge of the presence and degree of ITSC faults can be used by supervisory-level control to implement fault tolerant control. Results show that upon detection of a fault in the traction motor and its degree, the control takes appropriate steps, i.e., switching to the applicable traction input–output power efficiency map and gradually reducing the speed of the vehicle if it exceeds a predefined limit. These steps ensure that continued operation of the vehicle has a lesser risk of electrical system damage and fire than if the vehicle is allowed to continue on without any fault mitigation. The converse to continued, fault tolerant operation during an ITSC fault is to shut down the traction motor and render the vehicle immobile. However, there are safety issues associated with this as well, e.g., the motorist may be stranded in dangerous traffic or in a remote area. Further, we note the moving horizon observer is implicitly robust; however, in the presence of noisy measurements it may be necessary to increase the number of observer partitions and correspondingly decrease the partition time frame. In this case, because of the sinusoidal nature of the voltages and currents, the moving horizon observer will naturally “average” the estimated magnitudes. Future work will explore this aspect and other robustness aspects of the moving horizon observer. Additional future work includes performing control strategy verification experiments with a physically realized powertrain and reducing the computational time of the supervisory-level control and moving horizon observer. In part, the latter requires optimizing the number of horizons, horizon width, and the search algorithm. This is a dual formulation to the problem in model predictive control of determining optimal horizon parameters.

As computational power in vehicles continues to increase and processor prices decrease, we expect that using moving horizon observers for fault detection will become an increasingly attractive solution to improving electric machine safety, reliability, and repair costs.

The authors would also like to acknowledge the support of Greg Shaver and the Hoosier Heavy Hybrid Center of Excellence.

  • U.S. Department of Energy (Grant No. DE-EE0005568).

Appendix

In this section, details of the 2004 Toyota Prius ICE, battery pack, vehicle dynamics, mechanical power split device, maximum drive power, and electrical bus component models are described. Also, the conditions for ICE operation transitions from on and off are set forth.

Internal Combustion Engine.

The 2004 Toyota Prius powertrain has a 57 kW ICE with operating range between 1000 and 5000 rpm [22] as shown in Fig. 8. Broadening the power flow modeling ideas in Refs. [15,23], the power dynamics are represented by two first-order lag equations Display Formula

(A1)dPfueldt=1τfuelPfuel+1τfuelPfuelmax(ωice)ufuel
Display Formula
(A2)dPicedt=1τicePice+1τiceηice(Pice,ωice)Pfuel

Equation (A1) describes the fuel power, Pfuel, delivery dynamics, where Pfuelmax(ωice) is the maximum available fuel power for a given engine angular velocity ωice, ufuel ∈ [0, 1] modulates Pfuelmax, and τfuel = 0.04 s is the fuel delivery system lag [15,23]. In Eq. (A2), Pice is the engine mechanical output power, τice = 0.26 s is the average engine power lag due to combustion delay and crankshaft and flywheel inertias [15,23,24], and ηice(Pice, ωice) is the combined efficiency of the combustion and engine mechanical power delivery in which ηice(Pice, ωice)Pfuel(ωice) represents the indirectly commanded engine output power through the operation of the fuel delivery system. It follows that Display Formula

(A3)Pfuelmax(ωice)=Picemax(ωice)ηice(Picemax,ωice)

completes the specification of the variables. Values for Picemax and ηice are determined by least squares fits of appropriate engine data in Ref. [22]. Here, the functions Display Formula

(A4)Picemax(ωice)=α1ω¯ice+α0
Display Formula
(A5)ηice(Pice,ωice)=β20P¯ice2+β11P¯iceω¯ice+β10P¯ice+β01ω¯ice+β00

are sufficient, where ω¯ice=ωice/max(ωice) and P¯ice=Pice/max(Picemax) are normalized by their maximum values for numerical solution purposes. Table 1 lists the fit coefficients.

Further, the ICE angular velocity is expressed using Ωice=ωice2 by the conservation of power equation Display Formula

(A6)12JicedΩicedt=Pice,psd+Pice

Note Ωice=ωice2 defines a Lyapunov energy function Pice,psd is the soon to be developed power routed through the power split device, and Jice = 0.13 kg/m2 [25] is the estimated rotational inertia of the engine; ωice=+Ωice because the engine turns in only one direction.

Battery.

The Prius powertrain has a 21 kW discharge/charge power NiMH battery pack that provides a secondary power source and allows the capture and storage of excess vehicle kinetic energy via regenerative braking. The battery's SOC, W¯bat, dynamics are represented with a validated empirical formula [26], which has been modified [17] to include an additional quadratic power term to more accurately represent the 7.2 V 6.5 Ah six-cell NiMH modules [27,28] used herein Display Formula

(A7)dW¯batdt=ηbat(W¯bat,Pbat,vb)PbatnbWbatmaxηbat=ln(W¯bat+cb,1vb)+cb,2vbPbatnb+cb,3vb+cb,4vb(Pbatnb)2

In Eq. (A7), Pbat is the battery power input, Wbatmax is the battery's maximum rated storage energy; nb is the number of battery modules in the pack; vb = d, Pbat ≥ 0, for discharge; vb = c, Pbat ≤ 0, for charge; and cb,id/cb,ic,i={1,2,3,4} are discharge/charge coefficients obtained by the least-squares fit of instantaneous power efficiencies produced using NiMH module battery data in Refs. [27,28] and efficiency relationships in Ref. [29]. Table 2 lists the battery model parameters, which includes Pbatvb,bnd, the discharge and charge power bounds. Further, the Prius battery state of charge is restricted to W¯bat[0.4,0.8] [30].

Vehicle.

Vehicle dynamics have been described with a point-mass, linear motion dynamical model [15,17]. However, when this model has an input of power, a singularity occurs at zero velocity. To eliminate this singularity and develop a vehicle dynamics model consistent with our power flow approach, a Lyapunov energy function is defined, ϒv = v2, where v is velocity. Thus, the conservation of power to vehicle motion results in the power flow dynamical model Display Formula

(A8)12mvdϒvdt=Pd(v)+Prr(v,θr)+Pb(θr)+P2,L+Pwhl,psdPbrkPd(v)=v·(0.5ρairAfrCdv2sgn(v))Prr(v,θv)=v·(Crrmvgcos(θv)sgn(v))Pb(θv)=v·(mvgsin(θv))

where v=+ϒv considering only forward motion, Pd is the drag force power, Prr is the rolling resistance, Pb is the body force power due to gravity, Pwhl,psd is the wheel power from/to the power split device defined shortly, Pbrk is the frictional braking power, ρair is the ambient air density, mv is the total vehicle mass, Afr is the vehicle frontal area, Cd is the drag coefficient, Crr is the tire rolling resistance coefficient, and θv is the road grade angle. Table 3 lists the Prius vehicle parameters.

The frictional braking power, Pbrk, is equal to a maximum velocity-dependent braking power modulated by ubrk ∈ [0, 1]; the maximum braking power depends upon a smooth function that increases with velocity until a maximum of 50 kW braking power is achieved. Specifically Display Formula

(A9)Pbrk=Pbrkmax(ϒv)ubrk=50tanh(ϒv5)ubrk

Mechanical Power Split Device.

The PSD mechanically connects the ICE, generator electric drive (SPMSM1), and traction electric drive system (SPMSM2) via a planetary gear system, displayed in Fig. 9, to propel the vehicle, charge the battery, and start the ICE. Specifically, the planetary carrier is affixed to the output of the ICE; the SPMSM1 rotor is attached to the sun gear; and the ring gear is connected to the SPMSM2 rotor, which is linked (via additional gears) to the drive wheels. The planetary gear system dynamics are analyzed to develop the aforementioned Pice,psd and Pwhl,psd in Eqs. (A6) and (A8), respectively,

Display Formula

(A10)Pice,psd=(rs+rr)rs(P1,L|ω1,m|+ϵs)ωice
Display Formula
(A11)Pwhl,psd=rrγrs(P1,L|ω1,m|+ϵs)v

where rs is the sun gear radius, rr is the radius of the ring gear, ω1,m is the angular velocity of the sun gear and SPMSM1 (which are directly coupled), ϵs is a small constant ≪ 1 to prevent division by zero, and γ relates the vehicle velocity to the ring gear angular velocity, i.e., ωr = γv. To begin developing Eqs. (A10) and (A11), the dynamics of the PSD with the ring gear connected to the drive wheels are set forth in the context of Fig. 9, where the four planet gears are assumed to experience the same forces and thus can be replaced with a single-planet gear with combined mass and inertia Display Formula

(A12)Jsdω1,mdt=TsFpsrs
Display Formula
(A13)4Jpdωpdt=FprrpFpsrp
Display Formula
(A14)4mpdvpdt=FpsFpr+Fpc
Display Formula
(A15)Jcωcdt=(rs+rp)Fpc+Tc
Display Formula
(A16)mvdvdt=rrγFpr+Pd+Prr+Pb+P2,LPbrkv

The torque from/to SPMSM1 is denoted by Ts; Tc is the torque supplied by the ICE; Js, Jp, and Jc are the rotational inertias of the sun gear, planet gear, and planet carrier, respectively; ωp and vp are the planet gear angular and translational velocities, respectively; and Fps, Fpr, and Fpc are the forces at the interface of the sun and planet gears, ring and planet gears, and carrier and planet gear, respectively. Equation (A16) is based on knowledge that the force transferred to the ring gear is that which propels the vehicle (along with SPMSM2) and the assumption that the ring gear inertia is negligible compared to the overall vehicle. We note (i) rp = (rr − rs)/2, (ii) the planet carrier has the same angular velocity as the ICE, i.e., ωc = ωice, and (iii) the planet carrier inertia includes the carrier itself plus that of the ICE, since they are joined together and we assume that the planet inertia is small compared to that of the engine, and thus, the totality can be described with Jice. Further, we take Jp, mp, Js as negligible compared to the vehicle and ICE inertias. After applying the previous and algebraic manipulations, we obtain Display Formula

(A17)Jiceωicedt=rr+rsrsTs+Tc
Display Formula
(A18)mvdvdt=rrγrsTs+Pd+Prr+Pb+P2,LPbrkv

Next, we relate the sun gear torque to the SPMSM1 power applied using Ts=P1,L/(|ω1,m|+εs); P1,L is divided by the absolute value of ω1,m to obtain the expected response of the ICE and vehicle given that both P1,L and ω1,m can take positive and negative values. For example, (i) if P1,L is negative, power is being consumed by SPMSM1 and one expects the ICE speed to go down regardless of the sign of ω1,m (assuming Tc is constant), and (ii) if P1,L is positive, SPMSM1 is starting the ICE and its speed is expected to increase regardless of the sign of ω1,m. Finally, applying the expression for Ts, recognizing Tcωc = Pice, recasting Eqs. (A17) and (A18) into power flow equations and employing Lyapunov energy functions, Ωice=ωice2 and ϒv = v2, results in Display Formula

(A19)12JiceΩicedt=(rr+rs)rs(P1,L|ω1,m|+εs)ωice+Pice=Pice,psd+Pice
Display Formula
(A20)12mvϒvdt=rrγrs(P1,L|ω1,m|+εs)v+Pd+Prr+Pb+P2,LPbrk=Pwhl,psd+Pd+Prr+Pb+P2,LPbrk

verifying Pice,psd and Pwhl,psd in Eqs. (A10) and (A11), respectively.

Additionally, an expression for ω1,m is needed to determine Pice,psd and Pwhl,psd. Gears sharing a point of contact have the same tangential velocity at that point, thus ω1,m is available from Display Formula

(A21)rsω1,m+rrγv=(rr+rs)ωice

Table 4 displays the PSD parameters used herein. The PSD gear radii are given in terms of number of gear teeth rather than length, because the gear pitch is unknown and the pitch cancels out of the Pice,psd and Pwhl,psd expressions under the assumption all the gears have the same pitch.

SPMSM1 and SPMSM2.

Section 4 describes the SPMSM1 and SPMSM2 efficiency maps development for power flow modeling. This section provides the SPMSM1 and SPMSM2 efficiency map coefficients and maximum power expressions.

SPMSM1 Parameters, Efficiency Map Coefficients, and Maximum Power.

Table 5 lists the SPMSM1 parameters. The SPMSM1 efficiency fit coefficients are given in Table 6. Additionally, SPMSM1's maximum mechanical and electrical power are needed. The maximum SPMSM1 mechanical power during propelling and generating is modeled from given maximum torque versus angular velocity information with the maximum power modified by mildly extending the zero speed power to 1 kW, so engine start up is possible from zero speed and adding a curve segment centered at the torque region boundary at 60π rad/s to obtain a continuous first derivative function of ω1,m

Display Formula

(A22)P1,Lmax(ω1,m)={0.1547ω1,m+1,0ω1,mω1,ms11.061×107ω1,m30.1847ω1,m2+69.68ω1,m6.544×103,ω1,ms1<ω1,mω1,ms230.16,ω1,ms2<ω1,mω1,mu

where ω1,ms1=1798π/30rad/s,ω1,ms2=1802π/30rad/s, and ω1,mu=1000π/3rad/s. The maximum electrical power during propelling (generating is not needed in the supervisory-level model) is obtained from the efficiency and maximum mechanical power Display Formula

(A23)P1,invmax(ω1,m)=P1,Lmax(ω1,m)ηSPMSM1,prop(P1,Lmax(ω1,m),ω1,m)

SPMSM2 Parameters, Efficiency Map Coefficients, and Maximum Power.

Table 7 lists the SPMSM2 parameters. The SPMSM2 efficiency fit coefficients are given in Table 8. Further, expressions for SPMSM2's maximum mechanical and electrical power are needed for the no-fault and ITSC-faulted cases. When there is no fault, the maximum mechanical power during propelling and generating is modeled from given maximum torque versus angular velocity information where the maximum power is modified by mildly extending the zero speed power to 3 kW, so vehicle movement is possible from zero speed and adding a curve segment centered at the torque region boundary at 40π rad/s to obtain a continuous first derivative function of ω2,m

Display Formula

(A24)P2,Lmax(ω2,m)={0.3761ω2,m+3,0ω2,mω2,ms18.2869×105ω2,m30.4783ω2,m2+124.3ω2,m7.853×103,ω2,ms1<ω2,mω2,ms21.848×105ω2,m22.927×102ω2,m+53.65,ω2,ms2<ω2,mω2,mu

where ω2,ms1=1198π/30rad/s,ω2,ms2=1202π/30rad/s, and ω2,mu=200πrad/s. Maximum mechanical power data at each ITSC fault level are obtained from the control simulations in Ref. [3]. This maximum power data at each fault level are approximated using a cubic equation with coefficients determined via a least-squares data fit Display Formula

(A25)P2,Lmax(ω2,m,σ)=a(σ)+b(σ)ω2,m+c(σ)ω2,m2+d(σ)ω2,m3

where Table 9 lists the fit coefficients found for each fault level.

The maximum electrical power during propelling (generating is not needed in the supervisory-level model) is obtained from the efficiency and maximum mechanical power Display Formula

(A26)P2,invmax(ω2,m)=P2,Lmax(ω2,m)ηSPMSM2,prop(P2,Lmax(ω2,m),ω2,m)

Electrical Bus.

The battery pack, SPMSM1, and SPMSM2 electrical power flows come together in the DC–AC Inverter. Both of the drives' electrical power values include the inverter efficiency, which is assumed to also include any electrical bus losses. Thus, the electrical bus is taken as having loss-less power transfer efficiency, resulting in Display Formula

(A27)Pbat=P1,e+P2,e

where P1,e and P2,e are the SPMSM1 and SPMSM2 electrical power values that include inverter efficiency, respectively.

Interconnection Equations.

Interconnection equations are constraints for each mode that relate the states, algebraic variables, and continuous control inputs. First, the SPMSM1 electrical and mechanical connections are Display Formula

(A28)P1,ei={P1,emax(ω1,m)u1i,i=1η1,invηSPMSM1,gen(P1,Li,ω1,m)P1,Li,i=2,3,4
Display Formula
(A29)P1,Li={η1,invηSPMSM1,prop(P1,ei,ω1,m)P1,ei,i=1P1,Lmax(ω1,m)u1i,i=2,3,4

with ω1,m available from Eq. (A21) and inverter efficiency, η1,inv, of 0.98. Note that the maximum electrical power here is the inverter power divided by the inverter efficiency, i.e., P1,emax(·)=P1,invmax(·)/η1,inv.

Next, the SPMSM2 electrical and mechanical connections are Display Formula

(A30)P2,ei={P2,emax(ω2,m)u2i,i=1,2,3η2,invηSPMSM2,gen(P2,Li,ω2,m)P2,Li,i=4
Display Formula
(A31)P2,Li={η2,invηSMPSM2,prop(P2,ei,ω2,m)P2,ei,i=1,2,3P2,Lmax(ω2,m)u2i,i=4

where the inverter efficiency is η2,inv = 0.98. Note that the maximum electrical power here is the inverter power divided by the inverter efficiency, i.e., P2,emax(·)=P2,invmax(·)/η2,inv.

The electrical bus connections for each mode are Display Formula

(A32)Pbati=P1,ei+P2,ei

The power split device mechanical power connections are Display Formula

(A33)Pice,psdi=(rs+rr)rs(P1,Li|ω1,m|+εs)ωice,i=1,2,3,4
Display Formula
(A34)Pwhl,psdi=rrγrs(P1,Li|ω1,m|+εs)v,i=1,2,3,4

Finally, the frictional braking power interconnection equations are Display Formula

(A35)Pbrki=Pbrkmax(ϒv)ubrki,i=1,2,3,4

Engine Operation.

The Prius ICE may transition between off and on during vehicle operation. Practically, when the engine is off, only modes 1 and 4 are possible; modes 2 and 3 should be unavailable until the engine has finished start up and is on. Further, when the engine is off, no fuel usage or power output is expected, thus Display Formula

(A36)α2,α3=0
Display Formula
(A37)Pfuel,Pice,ufueli=0,i=1,2,3,4

The engine is turned off if all of the following conditions are satisfied:

  • EO1)Either the battery SOC has reached the nominal value Display Formula
    (A38)W¯bat(tp,0)W¯batnom=0.58
    or the reference velocity is decreasing at a rate below a threshold over a time interval Display Formula
    (A39)dvref(Δt)dt<aice-off=0.5m/s2
    where Δt ∈ [tp,0, tp,f].
  • EO2)The estimated power needed to meet the reference velocity is less than or equal to what can be supplied by SPMSM2 over a time interval Display Formula
    (A40)P2,Lmax(ωr,ref(Δt))P̂whl(Δt)
    where ωr,reft) = γvreft). The estimated power, P̂whl is assumed to be piecewise-constant over Δt. It is obtained from Eq. (A8) by making P̂whl=P2,L+Pwhl,psd and then solving for P̂whl using the shooting method such that the output of Eq. (A8) tracks the square of the reference velocity within a negligible error.
  • EO3)The estimated electrical power needed by SPMSM2 is less than or equal to what can be supplied by the battery alone over a time interval Display Formula
    (A41)P2,e(ωr,ref(Δt))Pbatd,bnd=21kW
    Consequent to the ICE off, Eq. (A37) is enforced; the power values immediately go to zero because the fuel from which power is derived is no longer input.

The engine is started if the reference velocity is nonzero and nondecreasing and any of the following conditions are met:

  • ES1)The battery SOC is below a threshold at tp,0Display Formula
    (A42)W¯bat(tp,0)<W¯bat,iceon=0.50
  • ES2)The estimated wheel power power needed to meet the reference velocity is greater than what can be supplied by the traction motor, SPMSM2, alone from tp,0 to tp,fDisplay Formula
    (A43)P2,Lmax(ωr,ref(Δt))<P̂whl(Δt)
  • ES3)The estimated electrical power needed by SPMSM2 exceeds that can be supplied by the battery alone over ΔtDisplay Formula
    (A44)P2,e(ωr,ref(Δt))>Pbatd,bnd=21kW

Upon the aforementioned conditions being met to start the engine, the ICE is driven up to speed by requiring that it have the minimum ICE operating speed of 100π/3 rad/s after a certain time, Δice,start = 0.25 s, i.e., Display Formula

(A45)Ωice(tice)(100π30)2,ticetp,0+Δice,start

and allowing ufueli,Pfuel,Pice0. Once the engine has started and reached minimum operating speed, mode 1 is no longer permitted, α1 = 0, until the engine is off again.

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Markel, T. , Brooker, A. , Hendricks, T. , Johnson, V. , Kelly, K. , Kramer, B. , O'Keefe, A. , Sprik, S. , and Wipke, K. , 2002, “ Advisor: A Systems Analysis Tool for Advanced Vehicle Modeling,” J. Power Sources, 110(2), pp. 255–266. [CrossRef]
Wiegman, H. L. N. , and Vandenput, A. J. A. , 1998, “ Battery State Control Techniques for Charge Sustaining Applications,” SAE Paper No. 981129.
Nam, E. , 2004, “ Advanced Technology Vehicle Modeling in PERE,” U.S. Environmental Protection Agency, Washington, DC, Report No. EPA420-D-04-002. http://nepis.epa.gov/Exe/ZyPURL.cgi?Dockey=P10004Y7.TXT
Meisel, J. , 2006, “ An Analytic Foundation for the Toyota Prius THS-II Powertrain With a Comparison to a Strong Parallel Hybrid-Electric Powertrain,” SAE Paper No. 2006-01-0666.
Staunton, R. , Ayers, C. , Marlino, L. , Chiasson, J. , and Burress, T. , 2006, “ Evaluation of 2004 Toyota Prius Hybrid Electric Drive System,” Oak Ridge National Laboratory, Oak Ridge, TN, Report No. ORNL/TM-2006/423. http://www.engr.uvic.ca/~mech459/Pub_References/890029.pdf
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Johnson, S. C. , Meyer, R. T. , DeCarlo, R. A. , and Pekarek, S. , 2016, “ Fault Detection in Surface PMSM with Applications to Heavy Hybrid Vehicles,” Purdue University, West Lafayette, IN, Technical Report No. 472. https://docs.lib.purdue.edu/ecetr/472
Campos-Delgado, D. , Espinoza-Trejo, D. , and Palacios, E. , 2008, “ Fault-Tolerantolerant Control in Variable Speed Drives: A Survey,” IET Electr. Power Appl., 2(2), pp. 121–134. [CrossRef]
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Diallo, D. , Benbouzid, M. E. H. , and Makouf, A. , 2004, “ A Fault-Tolerant Control Architecture for Induction Motor Drives in Automotive Applications,” IEEE Trans. Veh. Technol., 53(6), pp. 1847–1855. [CrossRef]
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Akrad, A. , Hilairet, M. , and Diallo, D. , 2011, “ Design of a Fault-Tolerant Controller Based on Observers for a PMSM Drive,” IEEE Trans. Ind. Electron., 58(4), pp. 1416–1427. [CrossRef]
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Xu, L. , Li, J. , Ouyang, M. , Hua, J. , and Li, X. , 2010, “ Active Fault Tolerance Control System of Fuel Cell Hybrid City Bus,” Int. J. Hydrogen Energy, 35(22), pp. 12510–12520. [CrossRef]
Meyer, R. , DeCarlo, R. , Meckl, P. , Doktorcik, C. , and Pekarek, S. , 2011, “ Hybrid Model Predictive Power Flow Control of a Fuel Cell-Battery Vehicle,” American Control Conference (ACC), San Francisco, CA, June 29–July 1, pp. 2725–2731.
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Johnson, V. , 2002, “ Battery Performance Models in ADVISOR,” J. Power Sources, 110(2), pp. 321–329. [CrossRef]
Markel, T. , Brooker, A. , Hendricks, T. , Johnson, V. , Kelly, K. , Kramer, B. , O'Keefe, A. , Sprik, S. , and Wipke, K. , 2002, “ Advisor: A Systems Analysis Tool for Advanced Vehicle Modeling,” J. Power Sources, 110(2), pp. 255–266. [CrossRef]
Wiegman, H. L. N. , and Vandenput, A. J. A. , 1998, “ Battery State Control Techniques for Charge Sustaining Applications,” SAE Paper No. 981129.
Nam, E. , 2004, “ Advanced Technology Vehicle Modeling in PERE,” U.S. Environmental Protection Agency, Washington, DC, Report No. EPA420-D-04-002. http://nepis.epa.gov/Exe/ZyPURL.cgi?Dockey=P10004Y7.TXT
Meisel, J. , 2006, “ An Analytic Foundation for the Toyota Prius THS-II Powertrain With a Comparison to a Strong Parallel Hybrid-Electric Powertrain,” SAE Paper No. 2006-01-0666.
Staunton, R. , Ayers, C. , Marlino, L. , Chiasson, J. , and Burress, T. , 2006, “ Evaluation of 2004 Toyota Prius Hybrid Electric Drive System,” Oak Ridge National Laboratory, Oak Ridge, TN, Report No. ORNL/TM-2006/423. http://www.engr.uvic.ca/~mech459/Pub_References/890029.pdf

Figures

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Fig. 1

The Toyota Prius hybrid powertrain architecture with power flows: (yellow) fuel power consumed, (green) electrical power produced, (red) electrical power consumed, (blue) mechanical power produced, and (orange) mechanical power consumed

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Fig. 2

The power envelope for SPMSM2 in propelling mode with degree of fault σ = 0.005, 0.01, 0.02, 0.05, and parameters in Table 7

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Fig. 3

Simulated Prius velocity tracking without fault: (solid line) simulated velocity and (dot) commanded velocity

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Fig. 4

Simulated Prius velocity tracking with ITSC faults of 0.5%, 1%, 2%, and 5%: (blue solid line) 0.5% fault, (red solid line) 1% fault, (green solid line) 2% fault, (cyan solid line) 5% fault, and (red dot) commanded velocity

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Fig. 5

Simulated Prius SPMSM1 (generator) and SPMSM2 (traction) mechanical powers with a 5% ITSC fault: (blue solid line) power and (gray dashed line) superimposed commanded velocity

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Fig. 6

Simulated Prius engine power for no fault and 5% fault level: (blue solid line) power and (gray dashed line) superimposed drive profile

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Fig. 7

Projected modes for the Prius simulation: (blue solid line) mode and (gray dashed line) superimposed commanded velocity

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Fig. 8

Prius ICE power output versus speed with fuel efficiency regions

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Fig. 9

Mechanical power split device sun-planet-ring gear system

Tables

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Table 1 Toyota Prius engine maximum output power and efficiency fit coefficients
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Table 2 Toyota Prius NiMH battery pack model parameters
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Table 3 Toyota Prius vehicle dynamics model parameters
Table Footer NoteaThe vehicle mass includes two 79 kg occupants.
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Table 4 Toyota Prius mechanical power split device model parameters
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Table 5 SPMSM1 parameters
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Table 6 SPMSM1 efficiency map fit coefficients: i = 1 for propelling, i = 2 for generating, and R2 is R-square error
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Table 7 SPMSM2 parameters
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Table 8 SPMSM2 efficiency map fit coefficients: R2 is R-square error
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Table 9 SPMSM2 maximum mechanical power curve fit coefficients

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