`
`: http:iitim.sagepubcomi
`
`
`
`integrated microprocessor control of a hybridi.c. engineibattery-electric automotive power train
`PW. Masding and J. R. Bumby
`Transactions of the institute of Measurement and Control 1990 12: 128
`DOI: 10.1177i014233129001200303
`
`The online version of this article can be found at:
`http:iitim.sagepub.comicontenii12i3i128
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`>> Version of Record — Jan 1, 1990
`
`What is This?
`
`1
`
`PAICE 2005
`
`Ford v. Paice & Abell
`IPR2014-00571
`
`1
`
`PAICE 2005
`Ford v. Paice & Abell
`IPR2014-00571
`
`
`
`Integrate :2 microprocessor control ol 8.
`hybrid LC. on 'ne/battery-electnc
`automotive power train
`
`by FEW. Masding, BSc, PhD and JR. Bumby, 35c, PhD, CEng, MIEE
`
`the development of a fully
`This paper describes
`integrated microprocessor control system for a hybrid
`i.c.-enginelbatrery-electric
`automotive power
`train.
`Torque control systems for the internal-combustion
`engine and the electric-traction motor are designed using
`digital transfer functions and indirect methods of tarqne
`measurement. Root-locus methods are used in all
`designs to provide first, critically dumped closed-loop
`response. In all cases simple proportional-plus-inlegral
`control proved sufficient to achieve this. An overall cycle
`speed controller allows the laboratory test system to be
`exercised over any test driving cycle and offers the ability
`to carry out sophisticated power sharing and transmis-
`sion shifting strategies.
`-
`
`Keywords: Hybrid vehicles, automotive power train
`control, microprocessor control, electric vehicles, i.c.-
`engines
`
`Symbols
`
`a
`fc
`f,
`
`Zero of gc(w')
`Counter value from flywheel speed probe
`Number of teeth on the flywheel speed probe
`gear
`Gain of gc(w')
`g
`gc(w') P+I Controller in w’—planc form
`3.:(1)
`Bilinear discretisation of gc(w')
`la
`Armature current, A
`if
`Field current, A
`J
`Flywheel inertia, kg in2
`K
`Constant relating dynamometer
`speed to load
`Flywheel count to roadspccd conversion
`factor
`
`K
`
`k,-
`Meg
`N
`Pm
`rf
`rw
`Ten,
`2}
`Tic
`t,
`T,
`
`go T,J2
`Equivalent vehicle mass, kg
`Speed, rev/min
`Inlet manifold depression, mbar
`Final drive ratio
`Vehicle wheel radius, to
`Motor torque, Nm
`Torque in gearbox output shaft, Nm
`Engine torque, Nm
`Controller design criteria, rise time, s
`Control system base sampling period {20 ms)
`
`School of Engineering and Applied Science, University of
`Durham, South Road, Durham DH1 SLE, England
`
`128
`
`4
`
`0'
`
`Engine throttle position, 0.9“ steps
`Demand throttle position, 09° steps
`Motor accelerator demand
`
`Controller design criteria damping factor
`Real part of closed—loop pole
`Damped frequency, radt’s
`
`1 . Introduction
`
`In this paper some of the control problems encoun-
`tered in designing and operating a ‘dnvc-byuwire’
`hybrid
`internal‘cornbustion
`(i.c.)
`cnginclbattcry-
`electric vehicle are examined. With two power sources
`in the drive: train, considerable flexibility in design and
`control of the complete system is possible. Various
`drive train arrangements have been investigated in
`previous computer-aided—dcsign studies (Willis and
`Radtke, 1985; Burke and Somuah, 1980) but most have
`favoured the parallel hybrid arrangement illustrated by
`Fig 1. This mechanical configuration consists of an
`i.c.-enginc and an electric traction motor connected
`mechanically in parallel so that both power sources are
`capable of driving the road wheels directly. The
`advantages of such a hybrid drive system stem from its
`versatility in being able to operate in pure electric mode
`in urban areas yet retaining an i.c.—cngine for high-
`spccd operation and long—range capability. By correct
`design, such a drive arrangement not only has the
`potential to reduce exhaust emissions in the urban
`environment substantially, but also of substituting up
`to 70% of the petroleum fuel used by the average road
`user (Forster and Bumby, 1988; Sandberg, 1980).
`Precisely how much petroleum substitution is achieved
`depends on the individual vehicle use pattern.
`To realise the full potential of the. hybrid drive,
`integrated control of both the prime movers and the
`common transmission is required. The problems asso—
`ciated with the development of such an integrated
`control system can be divided into two parts: mode
`selection; and component control. Mode selection is
`concerned with deciding whether the vehicle should
`run in an electric mode, an i.c.-cngine mode or whether
`the i.c.-cnginc and the electric motor should provide
`propulsion torque together. Selecting which of the
`many possible operational modes to use under given
`operating conditions is a complex problem and inter-
`acts strongly with the basic design of the hybrid power
`train. An optimisation study of these problems based
`on a computer simulation of different hybrid-vehicle
`
`Trans Inst MC Vol 12 No 3. 1990
`
`Downloaded from timsagepubcom at WAYNE STATE UNIVERSITY on January 21 . 2014
`
`2
`
`2
`
`
`
`Marding and Bumby
`
`free-wheel
`
`Le. englne
` Trunamlsslan
`
`
`
`
`
`
`Accelerator
`Brake
`
`__
`
`
`Motor
`Controller
`
`
`Fig 1
`
`Parallel hybrid electric vehicle drive train
`
`ratings and
`power-train configurations, component
`control
`strategies
`is discussed in some detail
`in
`Bumby and Forster (1987). The end result of the
`optimisation process
`is
`a mode controller which
`receives, as input, the driver’s brake and accelerator
`signals and then adjusts the torque demand to the
`engine and motor to meet the total demand. In addi-
`tion,
`it controls the gearbox, since selection of the
`correct gear ratio, to define engine/motor speed, has
`a critical effect on their efficiencies.
`
`Once the mode controller has decided on a gear
`ratio and torque demand to be met by each of the
`prime movers, it is necessary then to design individual
`components controilers which operate the engine,
`motor and gearbox so that they meet the appropriate
`demand as quickly as possible and, when necessary,
`also allow smooth transition between modes. Earlier
`work has examined the control problems relating to
`automation of discrete
`ratio transmission units
`(Masding 6! al, 1988). Conventional discrete ratio
`transmissions are ideal for this purpose since they offer
`the highest efficiency of any transmission system, and
`for this reason automation of such transmissions is
`attracting considerable attention (Main at at, 1937;
`Busca er a1, 1979).
`In this paper the additional component control
`problems relating to engine and motor torque control
`and smooth engine starting are addressed. One earlier
`system which tackled these problems was built in the
`USA by The General Electric Co. during the Electric
`and Hybrid Vehicle initiative and resulted in a
`microprocessor-controlled
`prototype
`hybrid
`car
`(Trummel and Burke, 1983; Somuah et a1, 1983). On
`the basis of preliminary design studies,
`this vehicle
`used an i.c.-engine and an electric traction motor con-
`nected mechanically in parallel. Control of the electric
`traction system was achieved by a chopper in the field
`
`Trans lnst MC Vol 12 No 8, 1990
`
`circuit and seriesr'parallel battery switching to vary the
`armature voltage. Starting resistances and clutch slip
`were thus neceSSary to move the vehicle from rest. In
`the present work, power electronic armature and field
`choppers are used to give smooth, efficient motor per-
`formance over
`the whole operating range and to
`remove the need for a clutch system. This same electric
`drive system has been successfully used in an opera-
`tional all-electric van produced by Lucas Chloride and
`Bedford (Manghan and Edwards, 1983). Bose at a!
`(1984) describe the control methods adopted in the
`HIV-1 but concentrate solely on transfer functiOns
`developed in the s-domain. In contrast,
`the control
`systems for the engine and motor presented in this
`paper make extensive use of digital models which have
`been previously developed to describe their dynamic
`characteristics (Masding and Bumby, 1990a; 1990b).
`Satisfactory performance of the completed con-
`trollers is confirmed by using an extensive laboratory
`test facility. The test facility is a full-scale version of
`a parallel hybrid drive train using a 35 kW i.c.~engine
`and a 32kW D.C. traction motor as prime movers.
`Both the engine and motor are coupled to a 4-Speed
`synchromesh gearbox via a toothed drive belt. To the
`rear of the gearbox a flywheel-and-dynamomcter com-
`bination provide a simulation of the loadings due to
`vehicle inertia and aerodynamic/tyre drag, respectively.
`Control of the laboratory system is carried out by an
`M68000 microprocessor system which is responsible for
`receiving data from the extensive range of transducers
`round the rig and responding with appropriate control
`signals to the throttle servosystem, power electronics
`and gearbox. This system allows the control algorithms
`developed in this paper to be fully tested under opera-
`tional conditions as well as in simulation. A complete
`description of the test bed facility is given in Bumby
`and Masding (1988).
`
`129
`
`3
`
`
`
`The final two modes are the regenerative braking
`mode and the accelerator ‘kick down’ mode. The latter
`provides the driver with full power from both the
`engine and the traction motor and is intended mainly
`for use in emergency conditions when all economy
`considerations are overridden. Finally,
`regenerative
`braking is used whenever the vehicle is braked, in order
`to recover some of the kinetic energy of the vehicle and
`return it to the batteries. Having the motor connected
`to the drive train permanently means that regenerative
`braking is always immediately available.
`Ail the above operating modes pose common control
`problems in that, after a particular mode has been
`chosen,
`it must be possible to schedule and control
`the torque output of both the engine and motor. In
`addition, to provide smooth transition between modes
`it
`is necessary to start and synchronise the engine
`with a moving drive train accurately. Torque schedul-
`ing is the responsibility of the overall vehicle-mode
`controller on the basis of a strategy arising from the
`optimisation study mentioned earlier; however, in this
`paper the secondary problem of individual component
`control
`to achieve the desired torques and to start
`the engine is addressed.
`
`3. Controller design
`
`Experience has shown that robust controllers suit-
`able for all the applications in the hybrid vehicle can
`be produced using proportional—plus-integral control.
`Such controllers can give satisfactory performance not
`only for torque control of both prime movers but,
`in addition, for engine speed on no-load and overall
`speed control through a cycle. An advantage of these
`low-order controllers is their speed of execution: during
`a typical driving cycle the main computer takes only
`3 ms to carry out cycle speed and prime—mover torque-
`control calculations. High Speed and accurate compu-
`tation is encouraged by the use of 32-bit
`integer
`arithmetic throughout. All controller design is carried
`out in the w’-plane using root-locus pole placement
`methods. Z-transfer functions are mapped into this
`plane by the transformation pair
`
`
`w =—
`2 (2—1)
`T, 2+1
`
`.
`
`"'(1)
`
`...(2)
`
`w’+21T,
`z=——---
`Md”:
`
`where T, is the sampling period. Owing to the similarity
`between the s and ill—planes,
`the proportional-plus-
`intcgral controller retains its usual form:
`
`...(3)
`
`W
`
`gc(W‘)=g(WfTa)
`
`Acceptable closed-loop performance is defined in terms
`of the rise time, t,, and the damping factor 6. These
`are defined for a second-order system by the equations:
`
`14" =0-de
`.1 __
`
`wd:fi?£'"§l
`
`...(4)
`
`.(5)
`
`Trans inst MC Vol 12 No 3. 1990
`
`Masding and Bomby
`
`TABLE 1: Possible operating modes for the parallel hybrid vehicte
`
`Mode
`
`Etectric mode
`
`l.C.-eogine mode
`
`Primary electric mode
`
`Primary i.c.-engine mode
`
`Hybrid mode
`
`Battery charge mode
`
`Regenerative braking
`
`Accelerator ‘kick—down'
`
`Description
`
`All propulsion power supplied by the
`electric traction system
`All propulsion power supplied by the
`i.c.-engine
`The electric traction system provides
`the principle torque, but when
`necessary its maximum torque is
`augmented by the
`engine
`The i.c.-engine provides the principal
`torque, but when necessary its
`maximum torque is augmented by the
`motor
`
`Both the i.c.—engine and the electric
`traction system provide torque Split
`between them in some way
`The i.c.—engine provides both the
`propulsion power and power to charge
`the batteries, with the traction motor
`acting as a generator
`During braking the vehicle kinetic
`energy is returned to the battery with
`the traction motor acting as a
`generator
`Essentially a primary i.c.~engine
`mode when full engine torque
`is allowed to give maximum
`acceleration
`
`2. Hybrid-vehicle control modes
`
`The different operating modes available with a
`hybrid drive system are summarised in Table 1. In
`general, the electric mode can be used in urban areas,
`for Short journeys and when the engine load would be
`small giving rise to low engine efficiency. It is always
`used for moving the vehicle away from rest, since a
`conventional clutch system is not included. thn the
`drive—train speed exceeds 1000 rev/min, the engine can
`be started and synchronised with the moving drive
`train to provide additional power if required. Such
`operation is possible owing to a free-wheel unit in the
`engine drive line which allows the engine to remain
`stationary when the rest of the drive train is in motion.
`Primary i.c.-engine mode is used when vehicle Speed
`and loading are both high, which gives high engine
`efficiency. When necessary, the engine torque can be
`augmented by the motor for rapid acceleration or hill
`climbing. Typically, the motor will be used to provide
`extra power if the engine output would otherwise
`exceed 90% of maximum, since this leads to ineffi-
`ciency. Over journeys with an exceptionally large
`amount of acceleration or hill climbing, the battery
`state of charge may become very low, but this can not
`be allowed to continue until
`the batteries are com—
`
`pletely depleted, since the vehicle would then be
`unable to move away from rest. To counter this
`problem, a negative torque may be scheduled from
`the motor so that the engine both drives the wheels
`and charges the traction batteries. As discussed in
`Bumby and Forster (1987), this mode is necessary but
`has low overall efficiency and so should be avoided if at
`all possible.
`
`130
`
`4
`
`
`
`Harding and Bumby
`
`
`Conn-alter Wlth Probe
`
`Engine No" Load Speed
`
`Throtlie
`
`Trnneter Function
`Gain Compensation
`
`S Ervo
`
`
`
`
`9din=0.0121Z-O(01137
`Nuisance“ -1.521z+1.senc
`
`
`
`
`3(2)
`0.045 (2—1)
`eTm=Jf
`
`
`Biz)
`22— 1.7914z+o.7955
`
`
`
` 2
`
`Engine
`Speed
`Demand
`
`Englne Speed
`Counter Value
`
`
`
` ”{2} Engine Speed (rpm)
`
`
`Fig 2 Engine speed control block diagram
`
`speed identified in Masding and Bumby (1990a) and
`repeated as
`
`AN(z) _ 0.838 — 1.5102-1 +1.922z-2
`138(2) _
`1 -— 1.7902-1 + 0.79524
`
`...(10)
`
`When this is connected to the required control algor~
`ithm and throttle servo-system, the block diagram of
`Fig 2 is produced.
`For large changes in throttle demand *- that is,
`greater than four steps per sample period — the throttle-
`position transfer functibn, G-Az),
`is non—linear as
`explained in Masding and Bumby (19903). However,
`for design purposes, small variations in throttle demand
`are assumed when 67(2) reduces to 1/2 producing
`a linear system which can be transformed to the w’
`plane for controller design. In order to produce an
`acceptably short synchronisation time for the engine,
`a system rise time of t,= 0.5 s and critical clamping are
`chosen as the design criteria. By Eons (4)—(6)
`this
`suggests closed—loop poles w'=4.71 1-14.71. Fig 3
`shows the compensated system root
`locus with the
`controller
`
`gc(w’) =0.012(
`
`w'+1.1)
`
`W!
`
`...(11)
`
`With this controller the presence of the closed—loop
`
`0'
`
`_4L
`_ tan (cos’1(—§))
`
`...(6)
`
`By Choosing a suitable rise time for a specific controller
`and adopting E = 0.707 for critical damping in all
`cases, the position of the required closed~loop poles
`is defined. These pole locations can only be used as
`an initial guide, however, because in reality the plant
`does not produce a second-order closed-loop system.
`Fine tuning of the controller design is achieved by an
`iterative process. Eqn (3) can be transformed back into
`the z-plane by the reverse mapping of Eqn (1) to give
`
`33(1):“ (3+ki)z+(ki“g)
`2—1
`
`“‘(7)
`
`where k: gan’Z. From this equation comes the
`discrete direct realisation for the controller output uk
`
`“1:: uk~1+(g+ki) €k+(k2“8) ek-l
`
`"(3)
`
`When referring to plant transfer functions for the
`control of both engine and motor torque, the coeffi—
`cients that are quoted apply to the following general
`discrete transfer function
`
`y(z) ._ b0+b1z‘1+... +b,.,,z”"1
`u(z) _ 1—alz‘1— . ..
`- anz‘”
`
`...(9)
`
`4. Engine starting and speed synchronisation
`
`Whenever the hybrid vehicle is operating in an
`all—electric mode or
`is stationary,
`the i.c.-engine
`will be uncoupled from the drive train by means of the
`one-way clutch. Since in either of these situations the
`engine is not required to provide torque,
`the most
`obvious strategy is to shut it down entirely in order
`to conserve petroleum fuel. Adopting this strategy
`means that the next time the engine is required it must
`be started and synchronised with the moving, and pos-
`sibly accelerating, drive train, before it can repiace or
`augment the torque supplied by the electric traction
`system. Consequently, a starting system is required
`which has fast response and no tendency to overshoot
`the prevailing drive-train speed, thus avoiding a shock
`torque in the drive shaft as the one—way clutch is
`engaged. Design of such a control system uses the
`transfer function relating throttle position to engine
`
`Trans inst MC Vol 12 No 3. 1990
`
`[nag s‘
`at
`
`til
`
`-ll
`
`-3a
`
`4.},
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`p
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`it
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`39
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`
`fwd-4.24354 +5 4.333199 -l.il6
`
`Fig 3 Compensated root locus for control of engine speed
`on no—iosd
`
`131
`
`5
`
`
`
`Masding and Bumby
`Enginl sound.
`N
`(rpm)
`seen
`Actual throttla
`
`
`
`Throttle pouit‘ion,
`(6.?“ scene!
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`and
`
`Simulated throttle
`position,
`8
`
`\xfiofual engine
`speed,
`N
`
`.‘
`
`V‘
`
`Simulated angina
`speed.
`N
`
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`
`14
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`:2
`
`:8
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`s
`
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`
`l
`.23
`
`l
`.AD
`
`.63
`
`.96
`
`5-2
`l.fl
`Time tsaes)
`
`Fig 4 Step test for the engine speed
`control system
`
`pole on the real axis modifies the system response so
`that the performance criteria are not achieved with
`exactly the calculated imaginary pole locations given
`above. The simulated and experimental closed-loop
`response of the system being shown in Fig 4. On this
`diagram the experimental
`throttle trace shows the
`step-rate non-linearity which was not included in the
`design. This is a consequence of the high gain needed to
`meet the fast system-response requirement, and the
`large errors present at
`the ' beginning of the step
`demand. Actual engine response is delayed by the
`throttlesstep-rate limitation causing it to lag behind the
`simulation; however, satisfactory damping is retained.
`
`4.1 Engine starting and load transfer
`
`When required, the warm engine will fire in, typic-
`ally, 250ms using the conventional electric starter
`motor, but there is a further delay while the engine
`accelerates up to the drive-train speed. Inertia starting
`used in the HTV-l project (Trummel and Burke, I933)
`allowad the engine to be completely coupled into the
`drive train in 300 ms, but the cost was the need for an
`additional clutch between the engine and the engine
`flyWhecl. A time analysis of the starting process is
`shown in Fig 5. In this experiment the meter was
`
`as illustrated by
`ioad,
`initially accelerating under
`the motor speed and torque traces. At time t: 0.45 s,
`however,
`the computer receives the start command:
`immediately it turns on the ignition and engages the
`starter motor. At
`the same time the throttle is
`opened 9“ and the computer then waits for the engine
`to fire. This is adjudged to happen when the engine
`speed passes 490 rev/min. Above this speed the starter
`motor is turned off and the speed control algorithm is
`entered to run the engine up to the drive-train speed.
`Synchronisation is deemed complete when the engine
`speed is within 45 rev/min of the drive-train speed
`which in this case is achieved within 0.7 s of the original
`command to start. At this stage,
`torque control
`is
`transferred to the engine which continues to accelerate
`the load. Total times for starting, speed synchronisa—
`tion and transfer of load are consistently about Is, as
`demonstrated by Fig 5. Starting a cold, and perhaps
`damp, engine is still an unreliable feature of modern
`cars; consequently the software must be ready to cope
`with failure to start. In the event of the engine faiiing to
`start after 5 s, the starter motor is disengaged, to allow
`battery recovery, before a second attempt is made.
`With a very hot engine, better starting is often achieved
`with full throttle opening and this might be a useful
`strategy for the computer to adopt on the second
`
`013.90 steps!
`Torque (Mm)
`a:
`
`
`leB
`Hotnr
`
`\.
`Speed \\\
`..
`f-“wq
`NJQV— 1
`"
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`Speed
`(rpm)
`1336
`
`lfiflfl
`
`Throttle position.
`
`123m “ifN‘w
`ream
`BBB
`
`Engine
`speed
`
`\r/tnroua
`'
`
`an
`
`35
`
`3B
`
`25
`
`2B
`
`15
`
`13
`
`5
`
`Initial ?°
`throttle
`
`hratklo
`position
`
`633
`
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`
`23a
`
`132
`
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`
`1.5
`
`1.5
`
`2.3
`Time iiflfiil
`
`z. ,
`
`Fig 5 Analysis 01 the engine starting
`and load transfer process
`
`Trans Inst MC Vol 12 No 3. 1990
`
`6
`
`
`
`attempt at starting it the engine had been operated very
`recently.
`In this work a warm engine has always been
`considered. The reasons for this are twofold: first the
`
`feasibility of the control algorithms were of principal
`importance and, if these would not work on a warm
`engine then cold-engine work is unnecessary; second
`much work on engine—management systems, including
`engine starting,
`is being carried out by the motor
`companies (Meyer et at, 1983).
`
`4.2 Fuel saving
`
`Although stopping the engine does save fuel, the cost
`is some delay in the availability of engine torque,
`however small. In addition, although having a station»
`ary engine saves fuel if the time between successive
`starts is sufficiently long, some fuel penalty must be
`associated with the starting process, making very short
`shutdown periods uneconomic. Potential fuel saving
`through stopping the engine under idling conditions
`prompted Volkswagen to incorporate this feature in its
`Fennel E range of cars, after which fuel savings of up
`to 30% over the otherwise conventional
`i.e.-engine
`vehicle have been reported (Schmidt, 1981).
`
`5. Prime-mover torque control
`5.1 Engine 1torque control
`
`responsible for
`is
`The engine torque controller
`producing from the engine
`the torque demand
`requested by the main vehicle controller. Torque is not
`measured directly for control purposes, but is calcu-
`lated from measurements of speed and inlet manifold
`depression. This step is necessary because the strain-
`gauge transducers, used to measure torque directly, are
`too expensive and unreliable for use in an operational
`vehicle. During the theoretical analysis of the engine,
`carried out in previous work (Masding and Bumby,
`19903),
`it was established that only one dynamic
`element is needed to describe the way that engine
`torque behaves when the input
`throttle angle is
`changed. This element, which is known as the manifold
`filling delay, can be described by a first-order digital
`transfer function which has speed-dependent coeffi-
`
`Masding and Bumby
`
`TABLE 2: Identified transfer-function coefficients and gain for the
`engine manifold filling delay
`
`Speed
`System
`
`revl‘min
`an
`on
`b,
`gain
`
`—43.05
`—10.26
`0.1 34
`0.786
`1000
`-3‘l .55
`710.964
`~0.177
`0.643
`1500
`—31.11
`~11.360
`0.153
`0.640
`2000
`—28.56
`- 9.985
`—€).185
`0.644-
`2500
`
`
`
`
`0.536 —€l.608 —11.i43000 —25.31
`
`cients. Identification experiments allowed the values of
`these coefficients to be determined at a number of
`engine operating speeds as set out in Table 2.
`A5 a consequence of this speed dependence, it might
`appear necessary to design a series of torque con—
`trollers, each valid over a small speed range, so that
`acceptable system performance is maintained at all
`times. Fortunately,
`the changes in engine gain and
`dynamics represented in Table 2 are not particularly
`great, and so it is possible to design one controller using
`the transfer function for 2000rev/min which achieves
`good results at all speeds. In general,
`the linearised
`torque-control system appears as shown in Fig 6. Once
`the appropriate gains and transfer-function coefficients
`have been added for operation at 2000 rev/min,
`the
`uncompensated root locus of the system, appears as in
`Fig 7a. This locus has two open—loop poles, one due to
`the manifold filling delay and one nearer the origin due
`to a digital filter used to process the manifold-pressure-
`transducer reading. In designing the controller the aim
`is to achieve a fast, critically damped response; Suitable
`design parameters for the engine are: rise time, t, = 0.2;
`and E = 0.707. By substituting these values into Eqns
`(4)—(6) the approximate pole locations are found to be
`w’ = — 11.78 $111.78. Using these locations as a guide
`the controller
`
`
`f
`
`aware—4C?)
`
`._.(12)
`
`was selected, which produces the modified root locus of
`Fig 7b. To test the completed system a step increase in
`demand of 10 Nm was applied with the results as shown
`in Fig 8a. As shown, the real system responded almost
`exactly as the simulation suggests it should. Away from
`
` Enghe Torque Demand
`+
` Manifold Fling Delay
`
`
`
`
` Mid l25=fifim(2l = bu z+h1
`
`413(2)
`z-a1
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`
`Tonya Model Gah
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`thtm= - (serswcuuz)
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`
`
`
`
`Fig 6 Block diagram of the linearised engine model
`
`Trans Inst MC Vol 12 No 3. 1990
`
`133
`
`7
`
`
`
`Mascling and Bumby
`
`has n“
`
`a
`
`Poles
`
`li-Blfitilg
`.
`
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`f”
`97.25
`Zeros liltt Show) /
`mm f
`
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`
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`“40“
`~39
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`9
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`Real H'
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`'
`dmmédmn
`fume
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`
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`i
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`
`Conpensatur
`
`gclu’ i=9. W WW
`
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`
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`
`~39
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`
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`E
`
`ill
`Real if
`
`Fig 7a Uncompensated locus forcontrol oi engine torque at
`2000 rev/min
`
`Fig 7b Compensated locus for centre! of engine torque at
`2000 rev/min
`
`Engine Torque. T..(N,p_:
`Throttle position.
`9
`(3.9" stens!
`‘3
`
`Engine Torque
`
`Inlet Manifold
`lane!)
`daprnssian. p-
`#30
`
`Torque demand
`
`hanifnld
`flIprossian,
`
`P-
`
`Thrufrln
`BUS:?:UH.
`
`H1“:
`Efldlnar‘mfifli
`
`.26
`
`-uo
`
`.63
`
`so
`.
`Tine (secs)
`
`«no
`
`use
`
`see
`
`253
`
`28m
`
`153
`
`tea
`
`m
`
`Fig 8a Simulated and experimental
`performance of the engine torque
`control system at 2000 rev/min
`
`he
`
`35
`
`JD
`
`2!
`
`2%
`
`IS
`
`16
`
`35
`
`so
`
`25
`29:
`
`in
`“a
`
`le)
`Englflfl Torque. TL.tN.n-i
`lain! manifold
`a...
`InhaFJ
`Throttle passriun,
`9 13.95 stuns)
`depreSSinn,
`A:
`annue
`335
`
`demand
`AB
`
`H=mndnl Fig 8b Simulated and experimental
`
`Engine torque
`T.-tN.o..I
`
`a
`__/—‘—E ~.
`H.
`r
`\\
`
`\
`
`Thruttla
`\‘xx‘e
`\rrj'” pasltiflrt.
`——'\
`
`a
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`depression. u.
`
`E=Expsriment
`
`.00
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`Tina i
`
`.56
`secs!
`
`250
`
`2m:
`
`15::
`
`tan
`so
`
`performance of the engine torque
`control system at 3000 rev/min
`
`the design speed of 2000 rev/min the gain and dynamics
`of the manifold-filling delay vary, causing some degre-
`dation of controller performance. At 3000rev/min, the
`reduced manifold-filling gain increases system rise time
`as illustrated by Fig St).
`A small change in demand of 10 Nm was chosen for
`initial testing so that the non-linear effects of throttle
`step rate would not affect the system. In practice the
`
`134
`
`control system must be able to cope with larger changes
`in demand. An example of when such a change does
`occur is the transition between acceleration and cruise
`in a driving cycle. Fig 9 shows that for a sudden 25 Nm
`drop in demand, which is representative of such a
`transition,
`the system response does not deteriorate
`significantly.
`In this diagram the torque-transducer
`trace is included to illustrate the accuracy of the derived
`
`Trans Inst MC Vol 12 N03,1990
`
`8
`
`
`
`Engine Torque
`(Mm)
`a5
`
`Indirect
`
`Masding and Bumby
`
`49
`
`35
`
`SB
`
`25
`
`23
`
`15
`
`Fig 9 Engine torque control system:
`comparison of indirect and direct torque
`measurements during a large step
`disturbance
`
`measurement, Tg‘lNYFN) \
`
`Demand signal
`
`
` Transducer measurement
`
`
`4.0
`
`6.!
`
`3.3
`
`to
`
`I2
`Time (secs)
`
`It
`
`torque measurement based on manifold pressure and
`speed. During the transient interval, direct comparison
`of the two torque readings is difficult since the torque
`transducer requires heavy filtering to remove noise
`effects and therefore demonstrates a delayed response.
`
`5.2 Electric motor torque control
`
`As explained in Masding and Bumby (1990b), the
`Lucas Chloride power electronics controller uses three
`control modes to cover the complete speed and loading
`range of the electric traction motor. The control unit
`contains both field and armature chopper circuits which
`allow it to provide full closed-loop control of both field
`and armature current. Each of the three control modes
`
`is
`is characterised by the way the field current
`controlled in response to changing brake or accelerator
`signals- In brief, relatively low loads and speeds give
`rise to the field-boost mode, whereby field current
`increases rapidly with accelerator demand. Once full
`
`rated field current is reached, then the full-field mode
`ensues with all control achieved via the armature. Full-
`field mode gives way to field-weakening mode at
`relatively high speeds. At speeds above ZOOOrevlmin
`there is a direct transition from field-boost mode to
`field—weakening mode, with rated field current not
`being reached.
`Each of the control modes results in a different
`transfer function relating accelerator demand, 9,” to
`torque output, Tam. In addition the gain and dynamics
`of the field-boost and field-weakening modes vary with
`the operating point defined in terms of initial speed,
`current values and accelerator setting.
`Fortunately,
`these variations
`in the gain and
`dynamics of the field-boost and field—weakening modes
`are not particularly great; hence it is not necessary to
`solve the problem of designing a controller which
`continuously adapts to the changing system. In fact it is
`possible to design a single fixed controller for each
`mode which maintains adequate performance over the
`
`Motor
`Torque
`D
`
`d +
`amen a
`
`Dla crate PH Controler
`
`Hm (I) = (gkaJZ'l-(kl—Qi
`at!)
`
`Motor
`3
`Power
`Electronics
`
`Eftecllve
`Vehlole
`Inerllu
`
`Electronic
`Flltar
`Buffer
`Amp.
`
`
`
`Current Control
`Mode
`Determlnnllon
`
`Motor Speed N rpm
`
`Torque Modal
`
`Temelirflfl Iflu
`
`Fig 10 Motor torque control system
`
`Trans Inst MC Vol 12 No 3.1990
`
`135
`
`9
`
`
`
`Mcsding and Bumby
`
`whole range of gains that are encountered. A simple
`software algorithm allows the computer to distinguish
`between the operating modes, and so the appropriate
`controller gains can be selected.
`The bloek diagram for the electric-motor torque-
`control loop is shown in Fig 10. As with the engine,
`indirect torque measurements, this time based on field
`and armature current, are used for control purposes. In
`this instance the pole-placement method is used to
`design a controller giving a critically damped response
`and a rise time of 150 ms. Three controllers are
`
`produced, each tuned to a transfer function representa-
`tive of an individual mode. Table 3 shows the three sets
`
`TABLE 3: Electric-motor transfer-function coefficients and torque-
`control parameters for all three operating modes
`
`Parameters Field boost
`
`Full field
`
`Field weakening
`
`a.
`32
`b1
`b2
`
`-1.359
`1.303
`1.316
`«4.424x 10-1
`—4.221 mo-t —4.269X10“1
`1.171 ><1'2
`1.521 xiii-2 —2.053><10‘a
`sesame"a —3.09?x10’°
`easeme—3
`
`Gain
`
`0.1946
`
`0.1019
`
`0.0553
`
`w'-plane g
`a
`
`3.0
`15.0
`
`z-plane g
`k,
`
`3.0
`0.45
`
`[nag 11'
`
`9.0
`15.0
`
`9.0
`1.35
`
`
`
`15.0
`9.0
`
`15.0
`1.35
`
`Files ll ~29.23
`42.24
`
`
`
`
`25 7—-
`
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`i
`I
`
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`4546.1?
`lllli
`
`
`Fig 11a Uncompensaied locus tor motor torque control in
`the fieldboost mode
`
`Real
`
`in“
`
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`11'
`
`Real
`
`Fig 111) Compensated locus for motor torque control in the
`field-boost mode
`
`136
`
`of model parameters and the corresponding controllers
`in both the w'-plane and the final z—plane form after
`bllinear discretisation. Part of the design is illustrated
`by Figs 11a and 11b which show the uncompensated
`and compensated root-loci for the field~boost transfer
`function.
`
`All three pole-placement designs have been tested by
`applying a step change of demand of lONm to the
`system. A full-scale simulation is car