oi
`
`
`
`Integrated microprocessor control of a hybrid i.c. enginelbattery-electric automotive power train
`P.W. Masding and JR. Bumby
`Transactions of the institute of Measurement and Control 1990 12: 128
`DOI: 10.1177i014233129001200303
`
`The oniine version of this article can be found at:
`http:/ftim.sagepub.comfcontenti'12/3i'128
`
`
`Published by:
`SAGE
`http:."fwww.sagepu blicationscom
`
`On behalf of:
`
` ‘
`
`..':;
`
`The Institute of Measurement and Control
`
`Additional services and information for Transactions of the Institute of Measurement and Control can be found at:
`
`Email Alerts: http://timsagepub.comi’cgi.'aIerts
`
`Subscriptions: http://tim.sagepub.com/subscriptions
`
`Reprints: hitp://www.sagepub_com/journaIsReprints.nav
`
`Permissions: http://www.sagepub.com;fiournalsPermissions.nav
`
`Citations: http://tim.sagepub.com;'content/12/3/128.refs.htmI
`
`>> Version of Record — Jan 1, 1990
`
`What is This?
`
`1
`
`PAICE 2215
`Ford v. Paice & Abell
`|PR2015-00792
`
` 1
`
`PAICE 2215
`Ford v. Paice & Abell
`IPR2015-00792
`
`

`
`lntegrato microprocessor control of a
`hybrid LC. on ‘no/batteryelectrlc
`automotive power train
`
`by P.W. Masding, Bsc, Phil and J.R. Bumby, Bsc, PhD, CEng, MIEE
`
`the development of a fully
`This paper describes
`integrated microprocessor control system for a hybrid
`i.cnengine/battery-electric
`automotive power
`train.
`Torque control systems for the intemal-combustion
`engine and the electric-traction motor are designed using
`digital transfer functions and indirect methods of torque
`measurement. Root-locus methods are used in all
`designs to provide fast, critically damped closed-loop
`response. In all cases simple proportional-plus-integral
`control proved suflicient 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
`jj
`
`Zero of gC(w')
`Counter value from flywheel spccd probe
`Number of teeth on the flywheel speed probe
`gcar
`Gain of g,,(w')
`g
`gc(w’) P+I Controller in w’—planc form
`g,,(z)
`Bilinoar discrctisation of gc(w')
`ta
`Armature current, A
`If
`Field current, A
`J
`Flywheel inertia, kg m2
`K
`Constant relating clynamometcr
`speed to load
`Flywheel count to roadspccd conversion
`factor
`
`K
`
`k,-
`Meg
`N
`Pm
`rf
`r,,,
`Tm
`Tf‘
`T,-C
`t,
`T,
`
`go T,/2
`Equivalent vehicle mass, kg
`Speed, rcvlmin
`Inlct manifold depression, mbar
`Final drive ratio
`Vehicle whccl radius, rn
`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, 0.9° steps
`Motor accelerator demand
`
`Controller design criteria damping factor
`Real part of closed-loop pole
`Damped frcqucncy, rad./s
`
`1 . Introduction
`
`In this paper some of the control problems encoun-
`tered in designing and operating a ‘drivc—by-wire’
`hybrid
`internabconibustion
`(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-aidcd—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.-engine 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.-engine 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.—cng'lnc mode or whether
`the i.c.—cnginc and the electric motor should provide
`propulsion torque together. Selecting which of the
`many possible operational modcs 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 BIC Vol 12 No 3. 1990
`
`Downloaded from tjm.sagepub.com at WAYNE STATE UNIVERSITY on January 21 , 2014
`
`2
`
`2
`
`

`
`Marding and Bumby
`
`free-wheel
`
`Trnnamlsalon
`
`Accelerator
`Elmka
`
`__
`
`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 controllers 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 at
`:11, 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 at at‘, 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 al, 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 3, 1990
`
`circuit and series/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 eiectronic 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 er al
`(1984) describe the control methods adopted in the
`HTV-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 confirrned 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 32I<W 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-dynamorneter 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 servosystern, 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
`
`

`
`Masding and Bumby
`
`TABLE 1: Possible operating modes for the parallel hybrid vehicle
`
`Mode
`
`Electric mode
`
`l.C.-engine 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.~engina
`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. When 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—whee1 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
`cfficicncy. 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. Ovcr 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
`
`The tinal 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.
`All 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—intcgral 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,
`z-t-1
`
`’
`
`w’-l~21T,
`z=——---~
`w,_2jT5
`
`mu)
`
`...(2)
`
`where T, is the sampling period. Owing to the similarity
`between the s and w-planes,
`the proportional~plus—
`integral controller retains its usual form:
`
`g.=(w‘)=g(wfTa)
`
`W
`
`...(3)
`
`Acceptable closed-loop performance is defined in terms
`of the rise time, t,, and the damping factor .5. These
`are defined for a second-order system by the equations:
`
`w’ =cr.iwd
`
`wd=E_-1,5-:3-3
`
`...(4)
`
`.15)
`
`Trans inst MC Vol 12 No 3, 1990
`
`4
`
`

`
`Controltar Wlth Probe
`Engme
`Engine No" Load Speed
`Speed
`Gain Compensation
`Trunster Function
`Demand
`9d(zJ=U.0121Z-0.01137
` N(?):.U.fl.’:l84:2-1.521z+1,92DU
`3(2)
`0.045 (2-1)
`
`
`Biz) z2~ 1.7e14z+o.7s5s
`
`Masding and Bttmby
`
`
`
`
` Ntzi Engine Speed (rpm)
`
`
`Fig 2 Engine speed control block diagram
`
`Cl’
`
`_ tan (cos’1(—§))
`
`...(6)
`
`speed identified in Masding and Bumby (1990a) and
`repeated as
`
`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
`
`set!) ='
`
`(E+kt)-°-'+(ki“8’)
`2-1
`
`mm)
`
`where k,-= gaT,./2. From this equation comes the
`discrete direct realisation for the controller output uk
`
`Ric“ ”k~1+(8‘i'kr) €k+(kz"“8) 5k-1
`
`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”"
`u(z) _ 1—a,z”1—- . ..
`-- a,,z'”
`
`...(9)
`
`4. Engine starting and speed synchronisation
`
`Whenever the hybrid vehicle is operating in an
`a1l—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
`
`AN(z) _ 0.838 — 1.5102-1 +1.922z'2
`138(2) _
`1 —- 1.7902-1 + 0.795;.--2
`
`...(1o)
`
`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 function, G-{(2),
`is non—linear as
`explained in Masding and Burnby (19903). However,
`for design purposes, small variations in throttle demand
`are assumed when G;-(z) 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 i'j4.71. Fig 3
`shows the compensated system root
`locus with the
`controller
`
`gc(w’) =0.012(
`
`w'+1.1)
`
`W!
`
`...(11)
`
`With this controlier the presence of the closed-loop
`
`may it’
`
`P
`
` .zi3n -4' 4. 33199
`
`Jan +
`-1.36792 ”
`
`4.333199
`
`Hg 3 Compensated root locus for control of engine speed
`on no—ioad
`
`Real u‘
`
`131
`
`
`
`Engine Speed
`Counter Value
`
`5
`
`

`
`Masding and Bumby
`Engina spend.
`N
`(rpm)
`Efififl
`
`
`Actual
`
`throttle
`
`Throttle pnnitiun,
`(fl.9° scene!
`[5
`
`B
`
`!&
`
`14
`
`12
`
`SH
`
`5
`
`I
`.23
`
`l
`.AD
`
`.63
`
`-96
`
`5.2
`1-H
`Time tapes)
`
`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 c1osed—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
`throttIe—step-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, 2S0ms 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-1 project (Trummel and Burke, 1983)
`allowed 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
`flywheel. A time analysis of the starting process is
`shown in Fig 5. In this experiment the motor was
`
`as illustrated by
`load,
`initially accelerating under
`the motor speed and torque traces. At time t: 0.45 s,
`however,
`the computer receives the start command:
`immediately it cums 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 failing 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
`
`Speed
`(rpm)
`1906
`
`lbflfl
`
`186$
`
`1235
`
`won
`BUB
`
`633
`
`#33
`
`ZBH
`
`132
`
`Throttle position.
`
`0 1B.9° stays!
`Torque (Rm)
`#5
`
`Engine
`spend
`
`Initial ?°
`throttle
`
`hratkln
`nusitian
`
`QB
`
`35
`
`3B
`
`25
`
`2B
`
`15
`
`1B
`
`5
`
`
`
`.53
`
`1.5
`
`1.5
`
`2.3
`Time (uses!
`
`z_ ,
`
`Fig 5 Analysis of the engine starting
`and load transfer process
`
`Trans Inst MC Vol 12 No 3. 1990
`
`3233
`
`255%
`
`24GB
`
`Zfififi
`
`moo
`
`izma
`BB3
`
`#33
`
`_JAg\_€///position,
`
`3
`
`Simulated throttle
`position,
`8
`
`\
`
`F
`sages,’
`denann/
`/
`
`/K // \\Acfual Engine
`‘
`speed.
`N
`.,
`
`V,
`
`simulated angina
`speed.
`N
`
` \
`
`6
`
`

`
`attempt at starting if 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 et1gine—rnanagernent systems, including
`engine starting,
`is being carried out by the motor
`companies (Meyer et al, 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
`Formel 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 torque 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 Burnby,
`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 coefficierrts and gain for the
`engine manifold lilting delay
`
`39935
`System
`
`revfmin
`an
`on
`b,
`gain
`
`-43.05
`-10.26
`0.1 34
`0.766
`1000
`-3‘! .55
`(10364
`-0.177
`0.643
`1500
`-31.11
`-11.3150
`0153
`0.540
`2000
`-28.56
`—— 9.985
`-0.185
`D.B44
`2500
`
`
`
`
`0.536 —€l.608 -11.143000 -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 rnanifolCl-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, i‘, = 0.2;
`and E = 0.707. By substituting these values into Eqns
`(4)—(6) the approximate pole locations are found to be
`w’ = — 1l.78ij11.78. Using these locations as a guide
`the controller
`
`g.(w’)=o.4(”’:'7)
`
`f
`
`__.(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
`
` E59516 Torque Demand
` Manifold Fling Delay
`
` Mid i35=4,5m(2l = bg z+h1
`113(2)
`Z‘i1
`
`+
`
`
`
`
`
`
`Tonya Model Gah
`
`Karim: - (s.srs+o.0s44r2)
`N
`
`Fig 6 Block diagram of the linearised engine model
`
`Trans inst MC Vol 12 No 3, 1990
`
`133
`
`7
`
`

`
`lnag ti’
`
`mi
`
`E»-
`
`1
`
`9
`
`391233935 ii talents
`
`mass
`
`
`
`compensator
`
`gclu’ l:@. 40:’ Fl}/ii‘
`
`
`
`
`"439
`
`H
`
`-29
`
`-13
`
`t
`
`ill
`Real if
`
`Mczsding and Bumby
`
`
`
`I
`
`_/
`H’
`
`Poles R-2&3
`‘
`
`97.25
`Zeros lllnt Shown)
`
`/r‘
`
`K
`
`X.
`
`""
`
`-“-
`
`-29
`
`-19
`
`a
`
`it
`Real M‘
`
`lnag rd’
`
`29
`
`-w
`
`l
`l
`
`“~39
`
`Fig Ya Uncornpensated locus forcontrol oi engine torque at
`2000 rev/min
`
`Hg Tb Compensated locus for centre! of engine torque at
`2000 rev/rnin
`
`Enqinu Torque. T..(N,p_:
`Throttle pnsirann.
`9
`(U.9° sreos!
`63
`
`Engine Torque
`
`Inlet mafltfuld
`lane!)
`depression. p,
`#30
`
`Torque dfimand
`
`nanifnld
`depression,
`
`;,_
`
`
`
`Thruerln
`B E) 5 x I : CI Fl .
`
`H
`9 —E E
`
`.20
`
`-uu
`
`5"’ " " " ’ ‘ "“' " ‘
`.63
`
`Time
`
`.Bfl
`(secs)
`
`«BB
`
`use
`
`EBB
`
`251!
`
`Zflfl
`
`153
`
`IEO
`
`are
`
`Hg Ba Simulated and experimental
`performance of the engine torque
`control system at 2000 rev/min
`
`lmm)
`Engine Tnraun. TL.tN.u-)
`Throttle position,
`9
`lB.9’ stuns)
`L!
`Tarcue
`demand
`
`-53
`
`lnlue manifold
`dnprassxnn,
`g_ lunar)
`335
`
`Engine torque
`r.-cN.u..=
`
`250
`
`zuu
`
`tan
`
`LED
`Sn
`
`Fig 8b Simulated and experimental
`performance of the engine torque
`control system at 3000 rewrnin
`
`E=Expsriment
`
`H=nndnl
`
`.26
`
`.hO
`
`.59
`
`.55
`Tina (secs!
`
`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 Sb.
`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
`
`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
`
`134
`
`Trans Inst MC ‘Jet 12 No 3, 1990
`
`56
`
`35
`
`JD
`
`2!
`
`29
`
`L5
`
`KB
`
`5 . a
`
`3:
`
`:9
`
`25
`
`L3
`lfl
`5.n
`
`
`
`j*——'\T..:,»
`r
`
`~.
`
`\
`
`\
`Thrnttla
`\‘E
`\‘rr:i'” pasitiurt.
`—’\
`\
`
`\\
`\:-H 3- depression. 0-
`
`a
`
`8
`
`

`
`Engine Torque
`(Rm)
`&5
`
`Indirect
`
`Masciing and Bumby
`
`measurement, T,,(NYpn) \
`
`
`
`Transducer measurement
`
`Demand signal
`
`
`
`¢.B
`
`6.!
`
`3.?
`
`IE
`
`[2
`Time (secs)
`
`14
`
`49
`
`35
`
`SB
`
`25
`
`2B
`
`15
`
`Fig 9 Engine torque control system:
`comparison of indirect and direct torque
`measurements during a large step
`disturbance
`
`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 Burnby (199%), 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
`
`Motor
`TOMUB
`Demand +
`
`Discrete P4-E Conrroler
`
`R
`
`rim (2) = (s'+k:)z+(k.-9!
`9(2)
`
`Current Control
`Mode
`Determination
`
`Torque Modal
`
`Temal-l1—{|1) Iflu
`
`
`
`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 2000rev;’min
`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, T,,,,. 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
`
`Molar
`8
`Power
`Electronics
`
`Eloclmnlc
`Filter
`Buffer
`Amp.
`
`Fig 10 Motor torque control system
`
`Trans Inst MC Vol 12 No 3.1990
`
`135
`
`9
`
`

`
`Mizsding 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 block 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,
`as
`b,
`b2
`
`-1.359
`1.303
`1.315
`«—4.424>< 10-‘
`-4.221 X10“ —4.269><10“‘
`1.171 x1‘?
`1.521 x1o-3 —2.05ax1o‘“
`a.5s9><1o"“ —3.o97x1o'°
`s.65ox1o"~"
`
`'
`
`Gain
`
`0.1946
`
`0.1019
`
`0.0553
`
`w’-plane g
`a
`
`3.0
`15.0
`
`z-plane g
`in
`
`3.0
`0.45
`
`[nag u’
`
`9.0
`15.0
`
`9.0
`1.35
`
`15.0
`9.0
`
`15.0
`1.35
`
`.-'
`
`-
`
`Piles ll ~29.23
`-l.l.2il
`
`as i-—-
`
`I,’
`Zems lllot Slmnl
`;"
`’-
`-646.1?
`I
`llilt
`
`
`-50
`
`-E5
`
`ll
`
`25
`
`full
`Real H‘
`
`Fig 11a Uncornpensated locus for motor torque control in
`the field-boost mode
`
`lnag u’
`
`L
`
`I."
`
`25
`
`~’’
`
`'4 '5 -
`I
`'
`'
`-tttliest ei 7.95132
`-26.89%
`
`,
`'5
`'3.
`8 :
`C3
`l
`
`'35[-
`
`-su
`
`'
`
`-25
`
`‘\-
`
`A
`
`ll
`
`gm’):il(u’li5)/H’
`
`25
`
`I
`
`fill
`Real H'
`
`Fig 111) Compensated locus for motor torque control in the
`fie-ld—boost mode
`
`136
`
`of model parameters and the corresponding controllers
`in both the w‘-plane and the final z-plane form after
`bilinear 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—p]acernent designs have been tested by
`applying a step change of demand of l0Nm to the
`system. A full-scale simulation is carried out simulta-
`neously with the step test in each case. Initial tests were
`carried out under similar operating conditions to those
`used to obtain the