Thermal Protection for AC Traction Motors Using Temperature Estimations to Calculate Torque Limits and Blower Speed Requirements

A portion of the disclosure of this patent document contains material which is subject to copyright protection. The copyright owner has no objection to the facsimile reproduction by anyone of the patent document or the patent disclosure, as it appears in the Patent and Trademark Office patent file or records, but otherwise reserves all copyrights whatsoever.

1. Field of the Invention

This invention relates generally to a control system for providing an AC traction motor with protection against damaging thermal overloads and, more particularly, to a system using information measured and supplied by a vehicle propulsion system controller to predict internal motor temperatures in near real time and to control the operation of the motor's inverter and blower fan.

2. Description of the Related Art

Alternating current (AC) traction motors are conventionally used in conjunction with electronic inverter drives as a means for transportation vehicle propulsion. The traction motor's capability to operate outside its continuous operating envelope for a short period of time is typically used to obtain optimum vehicle performance; however, the overheating which results reduces the motor life expectancy. Under these conditions, thermal protection is necessary to prevent damage of the stator winding insulation system and rotor cage.

Overload protection for rotating electrical machinery is intended to provide a means of preventing excessive overheating of the electrical winding system. Embedded winding sensors are used in conventional thermal protection schemes. Most large machines used in industrial applications generally employ resistance temperature detectors (RTDs) embedded in the stator windings for the purpose of directly sensing stator winding temperatures.

A more sophisticated technique is described in David R. Boothman et al., "Thermal Tracking--a Rational Approach to Motor Protection," IEEE Transactions on Power Apparatus and Systems, Vol. PAS-93 (Sept.-Oct. 1974), pp. 1335-1344. The Boothman article describes an overload analog relay with a design based on knowledge of the rotating machine's thermal circuit and feedback from an embedded winding detector. The system tracks motor temperature and provides a disconnect signal if a limit condition is exceeded.

Historically AC traction motors have not been protected against damaging thermal overloads. The location of the motors exposes temperature sensor lead cables to impacts from track debris and hostile environmental conditions. The possibility of false indications or damage to the sensor is usually considered a system reliability issue which precludes the use of embedded winding sensors.

Accordingly, an object of the present invention is to provide a method of supplying thermal protection for AC traction motors.

Another object of the present invention is to provide a method for determining and adjusting motor ventilation levels during use of AC traction motors.

The development of modern microprocessor-based controls enables the implementation of a computer algorithm in the vehicle control system which can help provide thermal protection automatically when necessary in a manner compatible with the vehicle power source and tractive effort characteristics. No abrupt change or sudden loss of tractive effort occurs. Another benefit of the control system is that it permits control of the motor ventilating blower speed as a function of motor temperature which assists in the minimization of vehicle on-board auxiliary power consumption.

Briefly, in accordance with a preferred embodiment of the invention, a thermal protection apparatus for an AC traction motor including a stator, a rotor, a blower fan, and an inverter comprises means for measuring ambient air temperature of the motor; means for measuring air atmospheric pressure of the motor; and means for using the measured ambient temperature and atmospheric pressure values for estimating a plurality of temperatures in the AC traction motor. Further provided are means for calculating a plurality of motor torque limits imposed by the plurality of estimated temperatures and means for controlling the inverter by providing a signal comprising the lowest of the calculated plurality of motor torque limits to an inverter controller. An optional fan control feature includes means for calculating a plurality of blower speed requirements imposed by the plurality of estimated temperatures and means for setting the blower speed of the motor to the highest of the calculated plurality of blower speed requirements.

The features of the invention believed to be novel are set forth with particularity in the appended claims. The invention itself, however, both as to organization and method of operation, together with further objects and advantages thereof, may best be understood by reference to the following description taken in conjunction with the accompanying drawings, where like numerals represent like components, in which:

FIG. 1 is a block diagram depicting the basic elements of a vehicle AC propulsion system;

FIG. 1a is a block diagram of a vehicle of the present invention;

FIG. 2 is a flowchart illustrating the steps in the thermal protection system of the present invention;

FIGS. 2a-2g include computer source code listings which may be utilized to control operation of a computer in accordance with the present thermal protection method;

FIGS. 2h-2m include a dictionary of parameters used in the computer source code of FIGS. 2a-2g;

FIGS. 3a-3b are graphs illustrating the expected responses of an AC traction motor to typical overload conditions and the effects of the thermal protection system.

FIG. 1 is a block diagram depicting elements of a vehicle AC propulsion system, and FIG. 1a is a block diagram of a vehicle of the present invention. A vehicle power source 10 may comprise any desired power source including, for example, a diesel engine/alternator, battery, or wayside power source such as a third rail or high voltage catenary. Electrical power is conditioned by solid state inverter drives shown as inverters 12 (only one of which is shown in FIG. 1) which have respective inverter controllers 16 to regulate the voltage and frequency applied to respective AC traction motors 14 in a manner consistent with the needs of the vehicle application. Inverters 12 are coordinated by a single propulsion system controller (PSC) 18 which controls each respective inverter controller and inverter for each axle (axles not shown) of a vehicle 510. A vehicle level control system 20 can be used to send commands to PSC 18.

FIG. 2 is a flowchart illustrating steps in the thermal protection system of the present invention. In one embodiment, the flowchart software resides within the propulsion system controller 18 and is executed by an on-board microprocessor. The effect on the motor propulsion system of the execution of the thermal protection system and the resulting inverter controller signals is a smooth and gradual reduction in motor torque, and therefore vehicle tractive effort, which simultaneously limits motor temperatures to predetermined values. No abrupt change or total loss of vehicle tractive effort occurs. The motor is not disconnected from its power source as is often the situation in conventional protection methods.

After the inverter is started, at a measurement step 108, the ventilating air ambient temperature (TAMB in .degree.C.) and the atmospheric pressure (PAMB in psi) of ventilating air are measured in the local area ambient in which the motor is operated by vehicle level sensors (512 and 514 in FIG. 1a). In one embodiment, ambient temperature is measured by a thermocouple in the vehicle, and ambient pressure is measured with a barometric pressure transducer in the vehicle.

At an input step 110, the measured ambient temperature and atmospheric pressure, as well as the following additional vehicle system input signals are supplied to the algorithm:

TQFB: motor air gap torque (from inverter control calculations) (Ib.sub.f -ft)

VL1: inverter DC link voltage (volts DC)

SLIPRPM: rotor slip speed (rpm)

SFBTM: traction motor speed feedback (rpm)

PWM: inverter mode (1 implies PWM mode)

DIESELSPD: engine speed (rpm)

BSPEED: blower speed status (0=stop, 0.5=half, 1=full)

IPHPK: fundamental phase current peak (amps)

RUNIV: inverter status Boolean (1 implies active inverter)

In a preferred embodiment, a plurality of temperature variables are stored and calculated for each of the stator, 518 rotor, 520 and inlet and discharge end regions 522 and 524, respectively (shown in FIG. 1a). In one embodiment, there are six stator winding nodes (T1SP, T2SP, . . . , T6SP) and five rotor core nodes (T1RP, T2RP, . . . , T5RP).

At step 112, the initial temperature variables are set and the phase current RMS value is calculated by dividing the input value of IPHPK (peak value of fundamental phase current) by the square root of 2.

Then it is determined whether the protection system must be initialized. An initialization is necessary when the microprocessor system is first activated and no values are present in the initial temperature variables. In the preferred system, even when the inverter and vehicle are shut down, a battery will keep the processor in the PSC active, and then each of the stator and rotor node temperature variables is set to the respective estimated temperature of the previous run. If initialization is needed, each of the stator and rotor node temperatures is set to an initial temperature value TINT which is calculated for each node using the equation:

wherein T.sub.p is the previous temperature stored immediately prior to inverter shut down; Tamb is the present ambient temperature; Tamb.sub.p is the previous ambient temperature; t is the elapsed time in minutes since PSC 18 has been without power; and const is the machine thermal time constant in minutes assuming a rotor at a standstill with no ventilation. The time constant, which varies with motor size, is 34.6 minutes in one embodiment.

An alternative method for determining motor initial conditions is to use the inverter to measure each motor's hot resistance immediately prior to providing power to the inverter. The hot resistance value can be used to calculate each motor's average winding temperature which is then used as a basis for the initial condition temperatures. This strategy has the advantage that it is independent of the status of blower 516 (shown in FIG. 1a).

At step 114 the RUNINV input value is used to determine whether the inverter is active (and thus a load on the system). If RUNINV is less than 1 (i.e., there is no load), then at step 116 each of the simulation losses CLOSS (no load core loss), W2 (fundamental secondary I R loss), FW (friction and windage loss), IPH (RMS value of motor phase current), LLOSS (stray load loss), WNS1, WNS2, WNS3 (stator winding harmonic losses), WNBAR (rotor bar harmonic loss), WNER1 (rotor end ring harmonic loss at the inlet end), and WNER2 (rotor end ring harmonic loss at the discharge end) is set to zero and SFBTM (traction motor speed feedback) is set to a finite value less than or equal to 0.1 (to prevent divisions by zero). Then the algorithm proceeds directly to step 130.

If RUNINV is 1, then at step 118 the synchronous RPM (RPMS) is calculated by adding the input signal magnitudes SLIPRPM (rotor slip speed) and SFBTM (which is set to a finite value such as 0.1 if less than a predetermined finite value). If RPMS is less than a predetermined value, then it is set to that predetermined value for the purpose of avoiding a zero division. Any small number less than or equal to 2 can be used as the predetermined value. Additionally at step 118 the SLIP is calculated by subtracting SFBTM/RPMS from 1, and the frequency (FREQ) is calculated by multiplying RPMS by the number of poles and dividing by 120.

The inverter mode can be either PWM (pulse width modulation) or a square wave. At step 120, the PWM input value is used to determine whether the inverter is in PWM or square wave mode.

If PWM is greater than 0 and the inverter is thus in PWM mode, at step 122 the phase voltage is calculated without harmonics using the equation of line 5926 in FIGS. 2a-2g and the stator, bar, and endring harmonic losses are set to 0. The equations shown in the algorithm of FIGS. 2a-2g are all for purposes of example only. There are a number of techniques for modeling the losses and the other relevant parameters, and the invention is not limited to the specific techniques shown in FIGS. 2a-2g.

If PWM is not greater than zero, then the inverter is in square wave mode. At step 124 the phase voltage is calculated with harmonics and the stator winding, rotor bar, and the rotor endring harmonic losses are calculated using the equations of lines 5934-5940 of FIGS. 2a-2g.

Regardless of whether the inverter is in PWM or square wave mode, at step 126 the fundamental secondary l.sup.2 R loss, the friction and windage loss, and the average stator slot eddy factor are calculated using the equations of lines 5943-5945 of FIGS. 2a-2g.

At step 128, no-load core loss and stray load loss are calculated using the equations of lines 5947-5948 of FIGS. 2a-2g.

Regardless of whether the inverter is powered down (RUNINV<1), at step 130, the individual stator, bar, and endring fundamental losses are calculated, and the combination of total harmonic and fundamental losses is calculated using the equations in lines 5950-5976 of FIGS. 2a-2g.

At step 132 air density adjustments due to altitude and ambient temperature variations are made to better estimate mass flow rates. First the equipment blower speed (SFBEB) is calculated by multiplying the motor's equipment blower constant by the engine speed input parameter and the blower speed input parameter. If SFBEB is less than a predetermined finite value such as 0.1 rpm, SFBEB is set to that predetermined value. The volumetric flow rates of the stator and rotor are calculated by multiplying SFBEB by the motor's stator core CFM (cubic feet per minute) fraction and rotor core CFM fraction, respectively. The air density calculations which are used to calculate the mass flow rates consider the prevailing ambient pressure and the rotor core duct and stator core duct air temperatures using the equations in lines 6004-6009 of FIGS. 2a-2g.

At step 134, the heat transfer coefficients are calculated. The heat transfer coefficients of the end windings and rotor end rings are assumed to be functions of rotor speed, and mass flow rate through the stator and rotor cores, as used in the equations of lines 6012-6015 of FIGS. 2a-2g.

At step 135, the rotor and stator duct air velocities (ft/min), Reynold's numbers, and duct heat transfer coefficients are calculated using the equations in lines 6017-6028 of FIGS. 2a-2g. The duct Reynold's Numbers are adjusted for air density changes.

At step 136, thermal conductances are calculated for the stator end coils and the rotor cage ends using the equations of lines 6105-6128. Stator, rotor, and total core air rise conductances are calculated by multiplying the specific heat with the respective mass flow rate as shown in lines 6135-6142 of FIGS. 2a-2g.

At step 138, the air temperature at the inlet and outlet ends of the motor, the stator and rotor nodal temperatures, the stator and rotor average temperatures, and the average stator and rotor resistances are calculated.

The temperature at the inlet end of the motor is calculated using the formula on line 6586 of FIGS. 2a-2g. To determine the temperature at the outlet end of the motor, the input value of BSPEED is evaluated. If BSPEED is less than 0.5, then the outlet air end air temperature is set to equal the inlet end air temperature because the blower is off. If BSPEED is not less than 0.5, then the air temperature at the outlet end of the motor is calculated with the total fundamental and harmonic losses, the total air rise conductance, and the ambient temperature, as shown in the equation of line 6590 of FIGS. 2a-2g.

In one embodiment, the stator and rotor nodal temperatures are calculated with numerical integration of system differential equations. For example, numerical integration methods such as Runge-Kutta methods or, as shown in lines 6595-6700 of FIGS. 2a-2g, Euler's method, can be used. The rotor and stator average temperatures and the rotor endring and rotor core differential temperatures are calculated using the equations in lines 6710-6714 of FIGS. 2a-2g. The average rotor and stator resistances per phase are calculated using motor specific constants and the calculated average rotor and stator temperatures.

At step 140, the blower speed requirements are estimated. To calculate stator winding blower speed control requirements, the input value of the blower speed (BSPEED) is evaluated. If BSPEED is less than 0.5, then it is determined whether the calculated stator coil temperature at a predetermined one of the nodes is greater than the HTOFS (stator winding blower transition temperature between 1/2 to full speed). If so, then the required blower speed of the stator (REQBSDS) is set to 1 (full speed). If not, then it is determined whether the calculated stator coil temperature is greater than ZTOHS (stator winding blower transition temperature between zero to 1/2 speed). If so, then REQBSDS is set to 0.5 (half speed). If not, then REQBSDS is set to 0 (zero).

If the input value of BSPEED was not less than 0.5, it is determined whether BSPEED is between 0 and 1. If so, then it is determined whether the calculated stator coil temperature is greater than HTOFS. If so, then REQBSPDS is set to 1. If not, then it is determined whether the calculated stator coil temperature is less than HTOZS (stator winding blower transition temperature between 1/2 to zero speed). If so, then REQBSPDS is set to 0. If not, REQBSPDS is set to 0.5.

If the input value of BSPEED was not less than 0.5 or between 0 and 1, then it is determined whether BSPEED is greater than 0.5. If so, than it is determined whether the calculated stator coil temperature is less than HTOZS. If so, then the REQBSPDS is set 0. If not, it is determined whether the calculated stator coil temperature is greater than FTOHS (stator winding blower transition temperature between full to 1/2 speed). If so, then REQBSPDS is set to 0.5. If not, REQBSPDS is set to 1.

To calculate rotor endring blower speed control requirements, the same steps as discussed respect to the calculation of stator winding blower speed control requirements are performed with the following different parameters:

outlet end endring in place of stator coil

HTOFER in place of HTOFS

ZTOHER in pace of ZTOHS

HTOZER in place of HTOZS

FTOHER in place of FTOHS

REQBSPDER in place of REQBSPDS.

To calculate rotor bar blower speed control requirements, the same steps as discussed with respect to the calculation of stator winding blower speed control requirements are performed with the following different parameters:

differential temperature between rotor end ring and rotor core in place of stator coil temperature

HTOFBAR in place of HTOFS

ZTOHBAR in pace of ZTOHS

HTOZBAR in place of HTOZS

FTOHBAR in place of FTOHS

REQBSPDBAR in place of REQBSPDS.

After the values of REQBSPDS, REQBSPDER, and REQBSPDBAR are obtained, at step 142 the largest value is determined for use as the BSPEED command in the motor. It is determined whether REQBSPDS is greater than or equal to REQBSPDER and REQBSPDBAR. If so, then EBSPCM (the blower speed command) is set to REQBSPDS. If not, it is determined whether REQBSPDER is greater than or equal to REQBSPDS and REQBSPDBAR. If so, then the EBSPCM command is set to REQBSPDER. If not, it is determined whether REQBSPDBAR is greater than or equal to REQBSPDER and REQBSPDS. If so, then the EBSPCM command is set to REQBSPDBAR. If not then the EBSPCM command is unchanged.

At step 144, the vehicle control level system uses the EBSPCM command to adjust the speed of the blower fan to the limiting case. This "adjustment" will result in a change of blower speed only if the latest EBSPCM command is different than the previous EBSPCM command.

Next, at step 146, the torque limits are estimated. Motor torque limits are calculated for each protected component by the following expression (the expression below is for the motor stator winding; other component expressions are identical in form):

wherein TQLIMS represents the motor toque limit imposed by the stator winding temperature; K1TQ represents the constant related to the maximum permissible stator winding temperature; K2TQ represents the constant related to desired time rate of change of motor torque; T2S represents the calculated stator winding temperature at a predetermined one of the stator nodes; and SGN (TQFB) represents a function which provides the sign (positive or negative) of the motor torque (provides for operation in motoring or braking modes).

The value of K1TQ is obtained using the desired value for K2TQ (rate of change of motor torque per degree of temperature) as follows:

wherein TQRATE represents the maximum permissible continuous value of motor torque with full air flow; and TSTALIM represents the maximum permissible continuous stator winding hot spot temperature corresponding to TQRATE.

The torque limit imposed by rotor endring temperature (TQLIMER) is calculated as discussed with respect to TQLIMS above with the following different parameters:

K1TQER in place of K1TQ

K2TQER in place of K2TQ

T3R in place of T2S

TERLIM in place of TSTALIM,

wherein T3R is the calculated rotor cage end ring temperature at the discharge end and TERLIM is the maximum permissible continuous rotor end ring temperature corresponding to TQRATE.

The torque limit imposed by rotor bar extension stress (TQLIMB) is calculated as discussed with respect to TQLIMS above with the following different parameters:

K1TQB in place of K1TQ

K2TQB in place of K2TQ

ABS(DELROT) in place of T2S

TBARLIM in place of TSTALIM,

wherein ABS(DELROT) is the absolute value of the calculated differential temperature between the estimated rotor cage endring temperature at the discharge end and a predetermined one of the estimated rotor core temperatures (discharge end) and TBARLIM is the maximum permissible differential temperature between rotor end ring and rotor core corresponding to TQRATE. Each of TQRATE, TSTALIM, TERLIM, and TBARLIM varies with motor model.

After the values of TQLIMS, TQLIMER, and TQLIMB are obtained, at step 148 the smallest value is used as the torque limit (TQLIM) command in the motor. It is determined whether ABS(TQLIMS) is less than or equal to ABS(TQLIMER) and ABS(TQLIMB). If so, then the TQLIM command is set to TQLIMS. If not, it is determined whether ABS(TQLIMER) is less than or equal to ABS(TQLIMS) and ABS(TQLIMB). If so, then the TQLIM command is set to TQLIMER. If not, it is determined whether ABS(TQLIMB) is less than or equal to ABS(TQLIMER) and ABS(TQLIMS). If so, then the TQLIM command is set to TQLIMB. If not then the TQLIM command is unchanged.

At step 150, the TQLIM command is used to adjust the torque call (TQFB) in the inverter control to the limiting case if the absolute value of the TQLIM command (the limit value imposed by the protection system) is less than the absolute value of TQFB (the torque call value requested by the vehicle level control system). The inverter control sends this new torque command value to the inverter which commands this new value of torque from the respective motor. When the motor is being derated by the protection system, the new torque limit value will be lower than the previous torque limit value to limit motor heating.

When PSC 18 is active, the protection system is executed continuously at regular intervals for each vehicle traction motor using the input signals defined above. At the conclusion of each interval, the following output signals are passed to vehicle system controller 20:

TQLIM: motor torque limit

ATEMPS: average stator temperature (.degree.C.)

ATEMPR: average rotor cage temperature (.degree.C.)

EBSPCM: motor ventilating blower speed command (0=stop, 0.5=one-half, 1=full)

R1AVE: stator winding resistance/phase (ohm)

R2AVE: rotor cage resistance/phase (ohm)

The traction motor torque limit (TQLIM), as calculated by the algorithm, will be an inverse function of the motor temperatures predicted by the algorithm. The relationship is shown in FIGS. 3a-3b.

R1AVE and R2AVE are supplied to the inverter controller for use in the control of the respective traction motor. The inverter controller uses the R1AVE and R2AVE values to adjust the supplied torque, frequency, and voltage of the motor to improve the motor control function accuracy.

Thus, temperature prediction is accomplished within the control system by numerical solution of a set of differential equations which characterize the transient thermal behavior of the motor. Input signals defined previously are used to determine the motor losses at any arbitrary operating condition in time. These losses along with the ambient condition of the ventilating air provide the thermal model with the necessary information required to perform the temperature predictions in near-real time while the vehicle is operating.

In addition to stator winding protection, the control system protects the motor rotor cage form thermal damage in two ways. Rotor cage end ring temperature is limited to guard against a reduction of ring mechanical properties which can occur at high temperatures, and the differential temperature between the rotor cage end ring and the bulk temperature of the rotor core is limited to prevent excessive fatigue damage of the rotor cage bar extensions.

FIGS. 3a-3b are graphs illustrating the response of an AC traction motor to a typical overload condition and the effects of the thermal protection function, with FIG. 3a representing a response for a cold motor start and FIG. 3b representing a response for a hot motor start. The graph of FIG. 3b includes a region wherein the motor is being derated by the protection system.

Torque limits are calculated for each motor subassembly for which protection is required. The worst case subassembly limit is selected as the governing value TQLIM. Higher temperatures result in lower values for the torque limits. The calculated torque limits vary as the overload progresses because the torque limit is continuously calculated during the operation of the invention's protection system. As some point in time the calculated torque limit and the value of torque requested by the vehicle controller can coincide. When this point is reached, the protection system overrides the vehicle torque request and the inverter controller begins to reduce traction motor torque, as shown in FIG. 3b.

Also shown in FIGS. 3a-3b is the status of the traction motor ventilating blower. Blower speed requests (Blower Speed Call) are calculated by the protection system for each protected subassembly within the motor. The limiting case value is selected as the governing value for the motor.

While only certain preferred features of the invention have been illustrated and described herein, many modifications and changes will occur to those skilled in the art. It is, therefore, to be understood that the appended claims are intended to cover all such modifications and changes as fall within the true spirit of the invention. ##SPC1##