Patent Grant
8625283
Exemplary embodiments of the invention relate generally to a system and method for improving the power handling capabilities of an electronic device, such as insulated gate bipolar transistor (IGBT) inverters. Moreover, such exemplary embodiments may relate to modeling, monitoring, and reducing the temperature of insulated gate bipolar transistor (IGBT) inverters.
Traction vehicles, such as, for example, locomotives, employ electric traction motors for driving wheels of the vehicles. In some of these vehicles, the motors are alternating current (AC) motors whose speed and power are controlled by varying the frequency and the voltage of AC electric power supplied to the field windings of the motors. Commonly, the electric power is supplied at some point in the vehicle system as DC power and is thereafter converted to AC power of controlled frequency and voltage amplitude by a circuit such an inverter, which includes a set of switches such as IGBTs. In some systems, the electric power may be derived from a bank of electrical batteries coupled to a leg of the inverter. The inverter may be configured to operate in a battery-charge mode and a battery-discharge mode. During the battery-charge mode, electrical energy from the field winding is used to charge the batteries. During the battery-discharge mode, electrical energy stored to the batteries is used to energize the field windings of the motors. The power handling capability of the inverter is limited, at least in part, by the ability of the IGBTs to dissipate the heat generated by the current in the IGBTs. Accordingly, it would be beneficial to have improved systems and methods for modeling the temperature of the IGBTs in the inverter. Improved temperature modeling techniques may be used to improve the power handling capability of inverters by improving heat dissipation. Improved temperature modeling techniques may also be used to provide techniques for monitoring IGBT temperature during operation.
Briefly, in accordance with an exemplary embodiment of the invention, there is provided an electronic device that includes a heatsink, a first dual IGBT coupled to the heatsink and configured to provide electrical power to a field exciter, a second dual IGBT coupled to the heatsink and configured to provide electrical power to a battery, a third dual IGBT coupled to the heatsink and common to the field exciter and the battery. The electronic device also includes a temperature sensor disposed in the heatsink, a cooling unit comprising a plenum and a variable source of air flow, and a controller. The controller is configured to receive a temperature reading from the temperature sensor and, based on the temperature reading, determine a desired level of cooling for at least one of the dual IGBTs, wherein an air flow rate provided by the cooling unit is determined based on the desired level of cooling.
Another exemplary embodiment provides a power system for a vehicle. The power system includes a double H-bridge comprising a heatsink, a first dual IGBT coupled to the heatsink and configured to provide electrical power to a field exciter, a second dual IGBT coupled to the heatsink configured to provide electrical power to a battery, and a third dual IGBT coupled to the heatsink and common to the field exciter and the battery charger. The power system also includes a temperature sensor disposed in the heatsink, a cooling unit comprising a plenum and a variable source of air flow, and a controller. The controller is configured to receive a temperature reading from the temperature sensor and, based on the temperature reading, determine a desired level of cooling for at least one of the dual IGBTs, wherein an air flow rate provided by the cooling unit is determined based on the desired level of cooling.
Another exemplary embodiment provides a method of cooling an electronic device. The method includes receiving a temperature reading from a temperature sensor disposed in a heatsink. The method also includes, based on the temperature reading, determining a desired level of cooling for at least one of a first dual IGBT coupled to the heatsink and configured to provide electrical power to a battery and a second dual IGBT coupled to the heatsink and common to a field exciter and the battery. The method also includes controlling an air flow rate of a cooling unit disposed in relation to the heatsink based on the desired level of cooling.
These and other features, aspects, and advantages of the invention will become better understood when the following detailed description is read with reference to the accompanying drawings in which like characters represent like parts throughout the drawings, wherein:
Since three legs 202, 204, and 204 corresponding to the three phases are used in the double H-bridge, the hardware of a three phase inverter is employed. The double H-bridge may be implemented in a single housing which uses a single heat sink to provide heat dissipation for the switches 104. In embodiments, the heat sink is cooled by forcing air over the heatsink. Due to double H-bridge topology, the power loss exhibited in each leg has a different power loss. Furthermore, the forced air cooling of the common heatsink can result in uneven cooling air flow about the three legs of the double H-bridge, making the thermal resistance related to each of the three phases non-uniform. The power handling capability of the double H-bridge will generally be limited by the hottest leg. Thus, the uneven power distribution and uneven cooling of the three phases may reduce the overall power handling capability of the double H-bridge. According to embodiments, a model for analyzing the thermal response of the double H-bridge is developed.
Thermal Impedance Models
Each dual module may include a pair of IGBTs, each IGBT coupled in parallel with its respective diode. As shown in
For each test configuration of
Considering the model described above, it is possible to determine the thermal effect that current in one of the phases has on the temperature under the hottest spot of each of the phases in the double H-bridge 200. Assuming that the current, Io, is applied to the dual IGBTs of phase B with both IGBTs switched on, the power dissipated by the pair of IGBTs can be computed according to the equation PB=Io*(VceB++VceB−). The temperature under the hottest spot of the dual IGBT of phase B due to the power dissipated by phase B is referred to as TB1. A temperature difference, δTB1, can be computed as TB1 minus the temperature of the air, Tair. If the current, Io, is applied to phase C the power dissipated by the phase C IGBTs can be computed according to the equation PC=Io*(VceC++VceC−) and the temperature at the hottest spot under phase B, TB 212, due to the power in phase C is referred to as TB2. Similarly, if the current, Io, is applied to phase A the power dissipated by the phase A IGBTs can be computed according to the equation PA=Io*(VceA++VceA−) and the temperature at the hottest spot under phase B, TB 212, due to the power in phase A is referred to as TB3.
The thermal resistances raising the temperature underneath phase B due to current in phases B, C, and A can then be calculated according to the following equations:
δTB1=RB*PB
δTB2=RBC*PC
δTB3=RBA*PA
In the above equations, RB is the thermal resistance raising the temperature underneath phase B due to the power in phase B, PB. RBC is the thermal resistance raising the temperature underneath phase B due to the power in phase C, PC. RBA is the thermal resistance raising the temperature underneath phase B due to the power in phase A, PA. Accordingly, the total temperature difference under phase B, δTB, can be computed according to the following equation:
δTB=RB*PB+RBC*PC+RBA*PA eq. 3.1
Repeating the same analysis for phase A and phase B yields:
δTC=RC*PC+RBC*PB+RCA* eq. 3.2
δTA=RA*PA+RBA*PC+RBA*PB eq. 3.3
In the above equations it is considered that RCB=RBC, RBA=RAB, and RCA=RAC. Further, thermal resistance may generally be expressed as the temperature difference divided by the power, as shown in the equation 3.4 below, wherein X can equal A, B, or C.
RXt=δTX/PX; where X=A,B or C eq. 3.4
Substituting equation 3.4 into equations 3.1, 3.2 and 3.3 yields:
RAt=δTA/PA=RA+RCA*(PC/PA)+RBA*(PB/PA) eq. 3.5
RBt=δTB/PB=RB+RBA*(PA/PB)+RBC*(PC/PB) eq. 3.6
RCt=δTC/PC=RC+RBC*(PB/PC)+RCA*(PA/PC) eq. 3.7
In the above equations, RAt represents an effective thermal resistance for phase A which if multiplied by the total power of phase A (PA) will result in the same δTA as the one in eq. 3.3 where the power through the three phases is different. Similar definitions apply for RBt and RCt. Using the equations described above, thermal tests can be conducted using the test configurations shown in
In the test configuration shown in
RAt_inv_TEST=δTA/Pphase=RA+RBA+RCA eq. 3.8
RBt_inv_TEST=δTB/Pphase=RB+RBA+RBC eq. 3.9
RCt_inv_TEST=δTC/Pphase=RC+RBC+RCA eq. 3.10
In the above equations, RAt_inv_TEST, RBt_inv_TEST, and RCt_inv_TEST are the thermal resistances, RAt, RBt, and RCt computed for the data collected using the test configuration shown in
In the test configuration shown in
RBt—hb—CB=δTB/Pphase=RB+RBA+RBC eq. 3.11
RCt—hb—CB=δTC/Pphase=RC+RBC+RCA eq. 3.12
In the above equations, RBt_hb_CB, and RCt_hb_CB are the thermal resistances, RBt and RCt computed for the data collected using the test configuration shown in
In the test configuration shown in
RAt—hb—CA=δTA/Pphase=RA+RBA+RBA eq. 3.13
RCt—hb—CA=δTC/Pphase=RC+RBC+RCA eq. 3.14
In the above equations, RAt_hb_CA, and RCt_hb_CA are the thermal resistances, RAt and RCt computed for the data collected using the test configuration shown in
Based on the test data described in Tables 1-4, it will be appreciated that the power in Phase A does not significantly affect the Phase B measurements, because RBt_inv_TEST is approximately equal to RBt_hb_CB. Similarly, the power in Phase B does not significantly affect the Phase A measurements, because RAt_inv_TEST is approximately equal to RAt_hb_CA. Therefore, RAB=RBA=0. Thus, equations 3.8 to 3.14 can be simplified to:
RAt_inv=δTA/Pphase=RA+RCA eq. 3.15
RBt_inv=δTB/Pphase=RB+RBC eq. 3.16
RCt_inv=δTC/Pphase=RC+RBC+RCA eq. 3.17
RBt—hb—CB=δTB/Pphase=RB+RBC eq. 3.18
RCt—hb—CB=δTC/Pphase=RC+RBC eq. 3.19
RAt—hb—CA=δTA/Pphase=RA+RCA eq. 3.20
RCt—hb—CA=δTC/Pphase=RC+RCA eq. 3.21
Using equations 3.15 to 3.21, the following equations 3.22 to 3.27 can be derived. Specifically, combining equations 3.17 and 3.19 provides:
RCt_inv−RCt—hb—BC=RCA eq.3.22
Combining equations 3.20 and 3.22 provides:
RAt—hb—CA−RAC=RA eq.3.23
Combining equations 3.21 and 3.22 provides:
RCt—hb—CA−RAC=RC eq.3.24
Combining equations 3.17 and 3.21 provides:
RCt_inv−RCt—hb—CA=RCB eq.3.25
Combining equations 3.18 and 3.25 provides:
RBt—hb—BC−RBC=RB eq.3.26
Also, for a validation check, equations 3.19 and 3.25 can be combined to provide:
RCt—hb—BC−RCB=RC eq.3.27
Equations 3.22 to 3.25 can be used to derive the parameters RA, RB, RC, RCB and RCA from the thermal test results. For each of the above thermal tests, a correction factor may be applied to the computed thermal resistances to account for the thermal grease 308 between the case 304 of the IGBT modules 302 and the heatsink 306 (
T_TEST=Tcase=Tair+Po*Rth—ch+PX*RXt
In the above formula, Rth_ch represents the case to heatsink thermal resistance and Po equals Pphase/2. Substituting 2*Po for PX and solving for T_TEST−Tair yields:
T_TEST−Tair=2*Po*[(Rth—ch/2)+RXt]
Thus,
[T_TEST−Tair]/Pphase=RXt_TEST=(Rth—ch/2)+RXt
As noted above in reference to
RXt=RXt_TEST−0.009 eq. 3.28
In equation 3.28, RXt_TEST can be determined using the following equation, where MaxTcaseX represents the maximum temperature taken from the thermocouples 500 (
RXt_TEST=(maxTcaseX−Tair)/(Vce1X+Vce2X)*Io eq. 3.29
The correction factor described above may be applied to the thermal resistances computed from the test data. A summary of those results are provided in Tables 5 and 6 below.
Table 5 show the thermal resistances computed from the test data with the correction factor applied. Applying equations 3.22 to 3.25 the values of table 5 yields the thermal resistances shown in Table 6. To validate the values shown in Table 6, the thermal resistances RCA, RCB, RC, RB, and RA may be used to compute estimated temperature readings for the test configuration shown in
In each of
In an embodiment, regression techniques may be used to derive equations for the thermal resistances RCA, RA, RC, RBC, and RB as a function of the flow rate of the cooling air. Test data can be collected for each of the test configurations shown in
From equations 3.22 to 3.77, the parameters used to calculate RA, RB, RC, RBC, and RCA are RCt_inv, RBt_hb_BC, RCt_hb_BC, RAt_hb_CA & RCt_hb_CA. The part-to-part variation of these parameters between different double H-bridges can be described using statistical analysis. For example, the data shown in tables 8, 10, 12, 14, 16, 18, and 20 can be input into a statistical modeling package, such as Minitab®. The statistical data for these parameters is shown below in table 21.
The statistical data can be used to determine the upper specification limits (USL) for each for each of the parameters RCt_inv, RBt_hb_BC, RCt_hb_BC, RAt_hb_CA & RCt_hb_CA and the upper specification limits for the resulting thermal resistances RA, RB, RC, RBC, and RCA. For example, using equations 3.22 to 3.27 and the mean and standard deviations computed for the thermal resistance parameters shown in table 21, a statistical analysis, such as a Monte Carlo analysis, can be applied to obtain the mean (μ) and standard deviation (σ) for RA, RB, RC, RBC, RCA at 200 SCFM. The mean and standard deviation for each thermal resistance RA, RB, RC, RBC, RCA at 200 SCFM can be used to compute the USL for each of the thermal resistances at 200 SCFM using the following equation:
Z=(USL−μ)/σ
In the equation above, Z represents the number of standard deviations that can fit between the upper specification limit and the mean value, and USL, μo, and σo represent the upper specification limit, mean, and standard deviation for a specific thermal resistance parameter RA, RB, RC, RBC, RCA at 200 SCFM. Using Z=3 and solving for the USL provides:
USL=σ*3+μ
Using a Z value of three ensures that the double H-bridge design will be robust enough to accommodate a large part to part variation. In table 21, the mean (μo) and standard deviation (σo) of each thermal resistance (RA, RCA, RC, etc.) have been identified for 200 SCFM cooling. Using these values and Z=3, USLRXX—200 SCFM can be identified. Then, the ratios of K1=μo/RXX200 SCFM, K2=USLRXX200 SCFM/RXX200 SCFM and K3=σo/RXX200 SCFM can be identified. Using these ratios, equations 3.22 to 3.27 and data from tables 7, 9, 11, 13 and 21 the USLRXX at all tested cooling conditions can be identified. An example calculation of the thermal resistance value RCA is shown below in tables 22 and 23. In this example, the statistical analysis for the thermal resistance RCA, using the data from table 21, provided a mean (μo) at 200 SCFM of 0.05092 and a standard deviation (σo) at 200 SCFM of 0.00153. These values were used in the example calculations shown below in tables 22 and 23.
Using the same method described above for each of the thermal resistances, RA, RB, RC, RBC, and RCA, provides the USL values shown below in table 24.
The USL values obtained for each thermal resistance, RA, RB, RC, RBC, and RCA, can then be used to derive regression equations for each of the thermal resistances. For example, regression techniques may be applied to the USL values to derive equations for computing the USL of each thermal resistance as a function of the air flow rate used to cool the heatsink. Applying regression techniques to the example data of table 24 provided the following regression equations:
RCA=−0.02328+0.30685/(1+((SCFM/2.216)^0.487)) eq. 3.30
RA=−0.05826+0.5357/(1+((SCFM/10.98)^0.46)) eq. 3.31
RC=−0.0145+0.394/(1+((SCFM/9.158)^0.568)) eq. 3.32
RBC=−0.01547+0.7537/(1+((SCFM/2.198)^0.779)) eq. 3.33
RB=0.045607+0.12515*exp(−SCFM/65.1)+0.291*exp(−SCFM/10.6) eq. 3.34
In an embodiment, thermal capacitances for each of the phases may be determined. To determine the thermal capacitances of each phase, thermal test temperatures may be obtained using the test configuration described in
Tau=RBt*CB
Applying the average RBt value at 150 SCFM (RBt_hb_BC_TEST−0.009) of 0.058868 degrees C./W and solving for CB yields:
CB=151/0.058868=2565 joules/Degree C.
Note that in the above equation, the value of RBt is not the USL value, but rather the measured test data as shown in table 17. Further, an equation describing the test cooling curve as a function of time may be expressed as follows:
deltaTB=(33.8−0.8)*exp(−t/151)+0.8
In the above formula, t is time, and deltaTB represents the change in temperature under phase B for a given time, t. Taken from the test data, 33.8 degr C. is the starting temperature at t=0 and 0.8 degr C. is the final temperature (offset) of the cooling curve. The formula is based on the assumption that the cooling curve has an exponential form. The equation above can be used to compute an estimated cooling curve that represents the estimates temperature of phase B, TB, minus the temperature of the inlet air, Tinlet, over time, t. The resulting curve can be compared to the measured cooling curve in order to prove its assumed exponential behavior, as shown in
Based on the above description, it will be appreciated that the thermal time constant, Tau, at a given air flow rate will be the same for each phase. Additionally, Tau may be determine according to the following formula, wherein Rth represents the thermal resistance and Cth represents the thermal capacitance:
Tau=Rth*Cth
Solving for the thermal capacitance, Cth, yields:
Cth=Tau/Rth
If the double H-bridge is operated with a different air flow rate, the thermal capacitance, Cth, of each phase will remain constant, but Tau and Rth will change. Thus, CB will equal 2565 J/degr C. for any air flow rate, but RBt will change from RBt(150 SCFM) and therefore Tau will change from 151 sec. For different phases at an air flow rate of 150 SCFM, it was shown that Tau remained 151 sec. Since RAt is different from RBt which is different from RCt, then CB will be different from CC which will different from CA. Solving for the phase C and phase A thermal capacitances, CC and CA, yields:
CC=Tau/RCt—hb—BC_TEST−0.009=151/0.049078=3077 J/degr C.
CA=Tau/RAt—hb—CA_TEST−0.009=151/0.065987=2288 J/degr C.
Using the thermal impedance models developed above, values can be determined for the thermal resistances and thermal capacitances applicable to each of the phases of the double H-bridge under various loading conditions and air flow rates. These values may then be used to predict the thermal behavior of the double H-bridge during normal operation. Being able to predict the thermal behavior of the double H-bridge during operation can enable a number of useful improvements to the double H-bridge, and associated control circuitry. For example, improved ventilation and overtemperature protection techniques may be developed, as described further below in reference to
Junction Temperature Estimation Models
Ipr_average=(Iload—av/n)+Imagn eq. 4.1
In the above equation, Ipr_average represents the average current in the primary winding of the transformer 804 or 810, n equals the turns ratio of the transformer, and Imagn represents the magnetizing current of the transformer 804 or 810. In an embodiment, n is approximately 2.875 for the transformer 810 corresponding to the battery 808 and n is approximately 6.33 for the transformer 804 corresponding to the field winding 802. Further, the magnetizing current, Imagn, may be approximately 30 amperes for both transformers 804 and 810. The average current in the primary winding of the transformer 804 or 810 is shown in
Furthermore, for a single period, T, the average current in the primary winding of the transformer, Ipr_average, will be divided between the two phases of the H-bridge, yielding I_phase 1_average, represented by line 1104, and I phase—2 average, represented by line 1106. Thus, the average current for a single phase over an entire period, T, will equal one half of Ipr_average, which is referred to a Ik and represented by line 1108. Further, the actual shape of the current waveform for a single phase is shown by lines 1108 and 1110, where line 1108 represents the current in the IGBT 104 and line 1110 represents the current in the diode 208. The current waveform for phase A and Phase C of the double H-bridge 200 is described further below, in reference to
a=di/dt=Vdc/[Lleak] eq. 4.2
b=di/dt=Vdc/[Lleak+Lmagn∥Lload*n2] eq. 4.3
In the above equations, Lleak represents the leakage inductance of the primary winding of the transformer 804 (approximately 29 uH) or 810 (approximately 23 uH), Lmagn is the magnetizing inductance of the transformer 804 (approximately 26 mH) or 810 (approximately 4.9 mH), Lload is the inductance of the load seen by the transformer 804 (approximately 0.22H) or 810 (approximately 1 mH), and n is the turns ratio of the transformer 804 or 810 (see
Based on the results for the rates, a and b, shown in tables 25 and 26 it can be appreciated that for all values of the link voltage, Vdc, 102 (
As shown in
IBave=Io*ton+Iod*[time that diode of other phases conduct] eq. 4.4
In the above equation, IBave is the average current through phase B, Io is the average of Ix & Iy, which is the average current in IGBTs in phase A or C during ton. Iod is the average current though the diode in phase A or C, during the time the diode is on. In both cases, this current also goes through the IGBT of phase B.
Since the falling rate di/dt of −b is fixed, there are three possible scenarios for the shape of the diode current. As used herein, t3 equals the half period, T/2, minus the time that the IGBT is on, ton. Further, tf (referred to by line 1308) is defined as the time that it would take for Iy (initial current of the diode) to diminish to zero, and equals Iy/b. The time t4 (not shown) is defined as the time during t3 that the diode carries current. Further, tz (not shown) is defined as the magnitude of the current in the diode at the time that the other IGBT 104 in the dual IGBT is switched on. The first scenario is shown in
Ipr—av_igbt=Io*ton*f eq. 4.5
For both phases A and C, the contribution of the diode current to Ipr_av may be determined according to the following formula:
Ipr—av_diode=Iod*tf*f eq. 4.6
The average current though the diode can be determining using the following equation:
Iod=(Iy+Iz)/2=Iy/2 eq. 4.7
Iod=(Iy+Iz)/2→Iz=2*Iod−Iy where Iz>0 eq. 4.8
In the scenario shown in
Ipr—av_diode=Iod*t3*f=[Iy−b*t3/2]*t3*f eq. 4.9
Based on the three scenarios described above, it can be appreciated that if tf is less than or equal to t3 then t4 equals tf. Further, if the Phase B IGBTs are switching OFF at zero current, there will be no switching OFF losses and the Phase A or phase C diodes have no Err losses.
From equation 4.1 above it can be appreciated that if the desired current, Iload_av, is known, the value Ipr_average can be calculated. Half of Ipr_average will come from one phase (50% on). Therefore,
Ipr—av/2=Ik=Ipr—av_igbt+Ipr—av_diode eq. 4.10
Further, the current, Iy, can also be expressed as a function of Io, as shown in the equation below:
Iy=Io+b*(ton/2) eq. 4.11
With regard to Ipr_av_diode, if tf is less than or equal to t3, equations 4.6 and 4.7 yield:
Ipr—av_diode=(Iy/2)*tf*f eq. 4.12
If tf is greater than, equation 4.9 yields:
Ipr—av_diode=(Iy−b*t3/2)*t3*f eq. 4.13
Since Iy is a function of Io and by definition Iy−b*tf=0, the following equation can be obtained:
tf=Iy/b=[Io+(ton/2)]/b
The equation above has two unknowns, Io and ton, so it cannot be solved in the form shown above. However, if tf>=t3 that indicates that ton is large enough, in combination with the level of Ix and the rate b, that there is not enough t3 time (T/2−ton) for the current through the diode to die off before the half period expires. This is clearly the case of low-voltage, high-current operation and t4=t3. On the other hand, if tf<t3 that indicates that ton is not large enough, in combination with the level of Ix and the rate b (and therefore Iy), that there is enough t3 time (T/2−ton) for the current through the diode to die off before the half period expires. This is clearly the case of high-voltage operation and t4=tf.
It will also be appreciated that for t4=t3 (case of tf>=t3) the calculation of Ipr_av_diode is very accurate. To resolve the issue of the two unknowns in identifying Ipr_av_diode (and from there Io), for t4=tf (case of tf<t3) the value, Iod, may be slightly overestimated, which will result in slightly overestimating Ipr_av_diode. By using t4=min (t3, tf) in the calculation of Iod=Iy−b*t4/2 it can be ensured that the duration of Iod is correct. So, the only overestimation is in estimating Io (and therefore level of Iod). By approximating t4=t3, the calculation of Ipr_av_diode will be very accurate when tf>=t3, and slightly overestimated when tf<t3. So, t4+t3 is used when estimating Io. This yields the following equations:
tf=t3=t4=T/2−ton=1/(2*f)−ton eq. 4.14
ton=(Vprim/Vdc)*0.5/fr eq. 4.15
Vprim=Vload*n eq. 4.16
In an exemplary embodiment, Vload_batt=80V, T/2= 1/1200 sec (fr=600 Hz) and Vload field may be computed according to the following equation:
Vload_field=0.161 Ohms*Ifield eq. 4.16a
Thus, knowing the levels of Ifield and Ibatt in the loads, equation 4.16a can be used to find the Vload_field or Vload_batt=80V may be used. Using these values through equation 4.15, ton can be determined for both battery and field excitation cases. Given that Iy=Io+b*ton/2 and also also Iy=Iod+b*t4/2 yields:
Iod=Io+(b/2)*(ton−t4) eq. 4.17
Using equation 4.14:
Iod=Io+(b/2)*[ton−1/(2*f)+ton]→Iod=Io−(b/2))*[(1/(2*f)−2*ton] eq. 4.18
From equations 4.5, 4.9, and 410:
0.5*Ipr—av=Ik=Io*ton*f+Iod*t4*f
Substituting tf from equations 4.18 and 4.14 yields:
0.5*Ipr—av=Ik=Io*ton*f+[Io−(b/2))*[(1/(2*f)−2*ton]*[1/(2*f)−ton]*f eq. 4.19
Referring back to equation 4.3, it is known that:
b=Vdc/[Lleak+Lmagn∥Lload*n2]
From equation 4.1:
Ipr_average=(Iload—av/n)+Imagn
Thus, equation 4.19 has only one unknown, Io. Manipulating the expression and solving for Io yields:
Ik=f*{[Io/(2*f)]−(b/2))*[(1/(2*f)−2*ton]*[(1/(2*f)−ton]}→Ik=Io/(2)−(b*f)*[(1/(2*f)−2*ton]*[(1/(2*f)−ton]→Io=2*Ik+b*f*[(1/(2*f)−2*ton]*[(1/(2*f)−ton] eq. 4.20
Battery Charging Example
Equation 4.2 can be used to determine values for Ix and Iy (
In table 27, Ibatt is the average battery current and Vdc is the link voltage 102. Additionally, the calculations shown in table 27 use a battery voltage, Vload_batt, of 80 Volts, frequency of 600 Hz, and a transformer turns ratio, n, of 2.875 for the transformer 810. Using these values, values for a and b were calculated as shown in table 27. Using the values for a and b shown in Table 27, the values shown in table 28 can be determined.
Based on the values from table 28, it will be noticed that as Vlink becomes higher, ton becomes smaller, and t3 becomes larger. Also, for the higher Vlink values, t3>tf and t4=t3. Thus, for these higher Vlink levels, Iz becomes zero, since the diode current dies off before the half period expires. Since t4=tf<t3 for the higher Vlink levels, t4+ton<half period=0.0008333 sec. Additionally, the peak Iy values, where the IGBT switches OFF, are large (284 A @ 1500V). To verify the accuracy of Io, ton, Iod, and t4, from the tables above, these values may be used to estimate the average load current (Isec_av), as shown below in table 29.
As discussed above, whenever t3>tf=t4 (cases of Vlink=1300V and 1500V above) Iod can be overestimated slightly. This results in a slight overestimation of the Ibatt=Isec_av, shown above in table 29. In all other cases (cases of Vlink 250V to close to 1300V) the estimations are very accurate.
Field Excitation Example
Equation 4.2 can be used to determine values for Ix and Iy (
In table 30, I_av_field is the average current in the field winding and Vdc is the link voltage 102. Additionally, the calculations shown in table 30 use a battery voltage, Vload_batt, of 80 Volts, frequency of 600 Hz, and a transformer turns ratio, n, of 6.33 for the transformer 804 (
Based on the values from tables 30 and 31, it will be noticed that, since Lb is large (25.63 mH), the rate b is small for all the operating range of Vlink. This can also be seen from the relative values of Ix, Io, Iy (close together). Because b is small, tf>t3 for all the operating range of Vlink. Thus, t4 is always greater than t3 unless the desired field current is too low, and therefore ton becomes very short. To verify the accuracy of Io, ton, Iod, and t4, from the tables above, these values may be used to estimate the average load current (Isec_av), as shown below in table 32.
As shown in table 32, since tf always larger than t3, there is no error in estimating Iod and, therefore, no error in estimating I_av_field. Using the values for Vbatt, Vdc (=Vlink), Ibatt (=I_av_batt) and If (=I_av_field) and using equations shown in tables 29 to 32, a computer model may be constructed to estimate values for ton_batt, Ipr_av_batt, ton_f and Ipr_av_f. The estimated values for ton_batt, Ipr_av_batt, ton_f and Ipr_av_f represent information known by the H-bridge controller, thus, the computer model may be used for non-real time estimations. Specifically, Vdc and the estimated values for ton_batt, Ipr_av_batt, ton_f and Ipr_av_f, may be used to estimate values for the phase current parameters Ix_B, Iss_B, Iz_B, Ix_batt, Iy_batt, Iz_batt, t4_batt, Id_batt (Ido), Iss_batt, Ix_f, Iy_f, Iz_f, t4_f, Id_f (Ido), and Iss_f, using equations derived above (and repeated in the Tables 28 to 32). The phase current parameters may then be used to determine power loss estimates for the IGBTs 104.
IGBT Pss PoA=IssA*Vce(IssA)
In the above equation, PoA is the power loss during ton, and PoA is zero during the rest of the period. Thus, for the full period average power:
PssA=IssA*Vce(IssA)*tonA*fr[Watts]
IGBT PswA: energy/pulse=[Eon(Ix—A)+Eoff(Iy—A)] and fr=pulses per sec→PswA=[Eon(Ix—A)+Eoff(Iy—A)]*fr[joules/sec=Watts]
The power loss for the phase A and Phase C diodes at Reverse Recovery, may be calculated from Iz using Err(Iz). During ON the diode losses may be calculated using parameters as function of Ido, where Ido={(Iz+Iy)/2} using Err(Ido). Using Phase A as an example:
Diode Pd=VfA(IdA)*IdA*(t4—A)*fr
Diode PrrA=ErrA(IzA)*fr
Ix—B=Ix—f+Ix_batt
At switching OFF the IGBT losses will be calculated from:
Iz—B=Iz—f+Iz_batt
During ON (steady state), the losses will be calculated from the average value of blocks 5, 6, 7 and 8 shown in
Iss—B=Ipr—av_batt+Ipr—av—f
Using the equation above, the switching off losses for the phase B IGBTs, IGBT Poff, may be computed using the following formula:
IGBT Poff=fr*EoffB(Iz—B)
The switching off losses for the phase B IGBTs, IGBT Pon, may be computed using the following formula:
IGBT Pon=fr*EonB(Ix—B)
The steady state losses (on-state) for the phase B IGBTs, IGBT Pss, may be computed using the following formula:
IGBT Pss=IssB*Vce(IssB)*0.5, where (0.5=(T/2)/T)
Furthermore, in phase B, each IGBT 104 is ON for the full half cycle. Thus, there is no current through the diodes of phaseB and, therefore, no losses associated with the diodes in phase B.
Double H-Bridge Optimization
Based on the equations described in relation to
The computer model for the full thermal behavior of the double H-bridge can be used to determine junction temperatures, Tj, of the IGBTs 104 based on any specifications. As an example, the specifications of EVOLUTION locomotives are shown in tables 33 and 34. In this particular example, it can be considered that the junction temperature, Tj, of the IGBTs 104 may be allowed to reach up to 130 Degr C. (BT used is Tj=150 degr C.) when operating at 49 Degr C. ambient (Tair=49 Degr C.+pre-heat from consist 5 Degr C.+pre-heat from blower/plenum 7° C.=61° C.). This will allow maximum thermal cycling of 130 Degr C.−61 Degr C.=69 Degr C. which will not restrict the long life of the device. Furthermore, for the present modeling, the H-bridge can be configured to provide a basis for comparing the improved double H-bridge of the present embodiments to a sub-optimal double H-bridge configuration. Specifically, the double H-bridge may be configured such that Phase A is used to power the battery 808 and Phase C is used to power the field winding 802. Using the thermal rating guidelines of table 33 as input, the computer model provides the junction temperatures, shown in table 35, for the sub-optimal double H-bridge design.
It can be appreciated from table 35 that for Vlink=Vdc=1500, the junction temperature, TjA, for the double H-Bridge exceeds the desired maximum temperature of 130 Degr C. Using the current limit (transient maximum conditions) of table 34 as input, the computer model provides the junction temperatures shown in table 36.
Based on the data above it can be seen that, if the ambient air temperature is high, the junction temperatures, TjA and TjB, for Vlink equal to 1300V or above, may exceed the desired junction temperature limit of 130 Degr C. In response to exceeding the junction temperature guideline of 130 Degr C., the double H-bridge controller may be configured to de-rate the current supplied to the load, as described further below in relation to
As shown in
As shown in table 37, by operating the battery charger in phase C and the field exciter in Phase A, the junction temperatures for all of the phases is below the 130 Degr C. junction temperature guideline. Furthermore, from the table 37 it can be seen that in the new double H-bridge design, TjA is always less than TjB and TjC. Thus, ventilation and thermal protection techniques used in the double H-bridge may be based only on phase B and phase C.
Estimating Junction Temperatures in a Double H-Bridge
TSair=dTS=dTS—B+dTS—C+dTS—A=PB*RSairB+PC*RSairC+PA*RSairA eq. 5.1
In the above equation, TSair represents the temperature difference between the temperature at the thermistor (sensor) position (TS) 1802 and the temperature of the cooling air (Tair), and PB*RSairB, PC*RSairC and PA*RSairA are the contributions of phase B, C, and A to the Sensor Temperature (TS) minus Tair. From equation 5.1, the value of TSair can be examined for different test configurations. In the test configuration shown in
TSair_inv=Pph*(RSairB+RSairC+RSairA)→TSair—inv/Pph=RSairB+RSairC+RSairA
In the above equation, TSair_inv represents the temperature at the sensor position 1802 minus Tair in the test with the configuration of
RSair_inv=RSairB+RSairC+RSairA eq. 5.2
In the test configuration shown in
TSair—AC=Pph*(RSairC+RSairA)→Tsair—AC/Pph=RSairC+RSairA
In the above equation, TSair_AC represents the temperature at the sensor position 1802 minus Tair in the test with the configuration of
Rsair—AC=RSairC+RSairA eq. 5.3
In the test configuration shown in
TSair—BC=Pph*(RSairC+RSairB)→TSair—BC/Pph=RSairC+RSairB
In the above equation, TSair_BC represents the temperature at the sensor position 1802 minus Tair in the test with the configuration of
RSair—BC=RSairC+RSairB eq. 5.4
Combining equations 5.2 to 5.4, the parameters for equation 5.1 can be determined and are shown below.
RSairB=RSair_inv−RSair—AC eq. 5.5
RSairA=RSair_inv−RSair—BC eq. 5.6
RSairC=RSair—BC−RSairB eq. 5.7
RSairC=RSair—AC−RSairA eq.5.8
For each of the test configurations shown in
RSair_config=(TS−Tair)/Pphase for this configuration
In the above equation, RSair_config is the thermal resistance between the temperature sensor and the ambient air for a particular test configuration. Exemplary RSair_config values for each test configuration, are shown below in tables 39-41.
Since RSair represents the thermal resistance between the temperature 1700 sensor and the cooling air, the thermal resistance between the case of the IGBT and the heatsink, Rth_ch, of the grease 308 is not a factor in computing the above values. Thus, the correction factor of 0.009° C./W is not subtracted from the values. Using the RSair values from tables 39-41 and applying equations 5.5 to 5.8, values for RSairB, RSairC, and RSairC1, and Rsair_A may be obtained, as shown below in table 42.
To verify the above method and results, the average power for each phase may be taken from the test data, in order to estimate TS-Tair (Est TS-Tair). The TS-Tair estimates may be compared with the test measured values of TS-Tair (Test_TS-Tair) that are based of the temperature sensor 1700, as shown below in table 43.
In addition to the three test configurations shown in
Using the above values for RSairB, RSairA, and RSairC1, estimated values for TS-Tair (Est TS-Tair) may be computed and compared to measured values for TS-Tair (Test_TS-Tair) based on temperature data gathered from the sensor 1700 for the test configuration of
Based on the data shown in tables 43 and 45, it will be appreciated that the method described herein provides an accurate prediction of the delta sensor Temperature (TS-Tair). Accordingly, the derived values for RSairB, RSairA, and RSairC may be used in determining the junction temperatures of the IGBTs based on the temperature sensor reading, as described further below. In an embodiment, Upper Specification Limits (USLs) may be derived for the thermal resistance values RSairB, RSairA, and RSairC. From equations 5.5, 5.6, and 5.7 it can be appreciated that the USLs for RSairB, RSairC and RSairA will depend on the USL's of RSair_inv, RSair_AC and RSair_BC. To determine the USL values for RSair_inv, RSair_AC and RSair_BC, values of RSair_inv, RSair_AC and RSair_BC were computed as discussed above, using six additional double H-bridge devices. The data gathered from these tests is shown below in tables 47, 49, and 51.
In tables 47, 49, and 51, the labels S1, S2, S3, S4, S5, and S6 represent the data gathered for the different double H-bridges used in the tests. The part-to-part variation of these parameters between different double H-bridges can be described using statistical analysis. For example, the data shown in tables 47, 49, and 51 can be input into a statistical modeling package, such as Minitab®, to obtain the mean (μ) and standard deviation (σ) of RSair_inv, RSair_AC and RSair_BC at an air flow rate of 200 SCFM. The statistical data for these parameters is shown below in table 52.
Using the statistical process outlined above in relation to tables 22 and 23, the mean and standard deviation for each RSair_config at 200 SCFM can be used to compute the corresponding USLs, using the following equation:
Z=(USL−μ)/σ
Using Z=3 and solving for the USL provides:
USL=σ*3+μ
An example calculation of the USL of RSair_inv is shown below in tables 53 and 54.
Using the same procedure for RSair_Ac and RSair_BC, the results shown below in table 55 were obtained.
The USLs for RSairB, RSairC, RSairA can be computed based on the USLs for RSair_inv, RSair_AC, and RSair_BC shown in tables 55 and using equations 5.5-5.7. From equation 5.5, the USL for RSairB can be determined, as shown below in table 56.
From equation 5.7, the USL for RSairB can be determined, as shown below in table 57.
From equation 5.6, the USL for RSairA can be determined, as shown below in table 58.
Applying regression techniques to the data shown in tables 56-58, regression equations that describe RSairA, RSairB, and RCairC as a function of air flow rate may be obtained. Applying curve fitting techniques to the data shown in table 56 yields:
RSairB=0.0115+0.3845*EXP(−SCFM/13.23)+0.066*EXP(−SCFM/78.6) eq. 5.9
Applying curve fitting techniques to the data shown in table 57 yields:
RSairC=6.47E−3+0.1406*EXP(−SCFM/16.23)+0.0257*EXP(−SCFM/139.8) eq. 5.10
Applying curve fitting techniques to the data shown in table 58 yields:
RSairA=7.14E−3+0.301*EXP(−SCFM/13.93)+0.044*EXP(−SCFM/83.67) eq. 5.11
In an embodiment, thermal capacitances between the temperature sensor position TS (1802) and the temperature of the cooling air (Tair) may be determined and are referred to herein as CSair_A, CSair_B, and CSair_C. First, from the average test data for 150 SCFM shown in table 58:
For the test configuration of
Po*ZSair—CA=Po*[RSairC∥(1/CCs)]+RSairA∥(1/CCs)
and Zsair_CA=Rsair_CA∥(1/CCAs), then:
If the time constants RSair_C*CSair_C=RSair_A*CSair_A are equal to τ0, then:
From the above, since RSair_C+RSair_A=0.0236763+0.0047235=0.02839998=RSair_CA, it can be confirmed that the time constant, τ0, is the same for Rsair_CA, RSair_C, and RSair_A. Similarly, for the test configuration of
To test the assumption that τ0 is the same for RSair_inv, RSair_C, RSair_B, and Rsair_A, the thermal capacitances for C and A powered, B and C powered and B, C and A (inverter) powered can be determined by collecting test data for each of the test configurations shown in
τ_inv=196 sec
τ—BC=190 sec
τ—CA=186 sec
The value TS_XX−Tinl may be estimated using the following equation:
TS—XX−Tinl=(starting temperature−ending temperature)*exp(−t/τ)+ending temperature
The estimated value for TS_XX−Tinl may then be compared it with the test data, as shown in
Using the average of the thermal time constants shown above, (196 sec, 190 sec, 186 sec) provides:
τ_inv=τ—BC=τ—CA=τ—A=τ—B=τ—C=190 sec
And taking into account that t=Rth*Cth, the thermal capacitances can be calculated using the average test data for 150 SCFM from table 59, as shown below:
CSair—B=190/0.0131241→CSair—B=14,477 J/degr C. eq. 5.12
CSair—A=190/0.0047253→CSair—A=40,209 J/degr C. eq. 5.13
CSair—C=190/0.0236763→CSair—C=8,025 J/degr C. eq. 5.14
Based on the above data, it will be appreciated that the affect of the thermal capacity of phase A on the change of temperature of the Sensor is much weaker than the affect of the thermal capacity from phases B and C, since the thermistor is situated between phases B and C.
The thermal resistances and thermal capacitances derived above can be used to determine thermal impedances for ZSairA, ZSairB, and ZSairC. In an embodiment, the thermal impedances may be used to generate a computer model for determining the junction temperatures of the IGBTs 104 based on the reading from the temperature sensor.
To determine the junction temperatures, the temperature difference between the temperature sensor 1700 and each phase's case may be determined. As discussed above, TA=heatsink temperature hot spot under device in phase A, TB=heatsink temperature hot spot under device in phase B, and TC=heatsink temperature hot spot under device in phase C. TA, TB, and TC can be determined according to the equations 3.1, 3.2, and 3.3 using RCA=RAC=0. Accordingly:
TA=PA*RA+PC*RAC+Tair
TB=PB*RB+PC*RBC+Tair
TC=PC*RC+PB*RBC+PA*RCA+Tair
In the above equations, PA, PB, PC are the power loss through both IGBTs and diodes in phase A, B, C respectively. Furthermore, the thermal resistance parameters RA, RB, RC, RCA, and RCB may be determined based on the air flow rate, using equations 3.30 to 3.34. The summary of the USLs for these parameters is shown in table 24.
Equations for TA, TB, and TC may be derived using Tsensor. The values for TA, TB, and TC derived using Tsensor are referred to herein as TAS, TBS and TCS, respectively. Based on the description provided herein, it is known that:
TB=TSair+Tair+TBS=PB*RB+PC*RBC+Tair
TSair=RSairA*PA+RSairB*PB+RSairC*PC
Combining these equations yields:
TBS=(RB−RSairB)*PB+(RBC−RSairC)*PC−RSairA*PA
The contribution of PB to phase B may be expressed as:
RB−RSairB=RB—BS eq. 5.15
The contribution of PC to phase B from phase C may be expressed as:
RBC−RSairC=RC—BCS eq. 5.16
Thus, the equation for TBS may be expressed as:
TBS=RB—BS*PB+RC—BCS*PC−RSairA*PA eq. 5.17
Similarly, with regard to TCS, it is known, based on the description provided herein, that:
TC=TSair+Tair+TCS==PC*RC+PB*RCB+PA*RCA+Tair
Thus, TCS becomes:
TCS=(RCB−RSairB)*PB+(RC−RSairC)*PC+(RCA−RSairA)*PA and if
The contribution of PB to phase C from phase B may be expressed as:
(RCB−RSairB)=RB—CBS eq. 5.18
The contribution of PC to phase C may be expressed as:
(RC−RSairC)=RC—CS eq. 5.19
The contribution of PA to phase C from phase A may be expressed as:
(RCA−RSairA)=RA—CAS eq. 5.20
Thus, the equation for TBS may be expressed as:
TCS=RB—CBS*PB+RC—CS*PC+RA—CAS*PA eq. 5.21
Similarly, with regard to TAS, it is known, based on the description provided herein, that:
TA=TSair+Tair+TAS==PA*RA+PC*RAC+Tair
TSair=RSairA*PA+RSairB*PB+RSairC*PC
Combining these equations yields:
TAS=(RA−RSairA)*PB+(RBC−RSairC)*PC−RSairB*PB
The contribution of PA to phaseA may be expressed as:
(RA−RSairA)=RA—AS eq. 5.22
The contribution of PC to phaseA from phase C may be expressed as:
(RBC−RSairC)=RA—ACS eq. 5.23
Thus, the equation for TAS may be expressed as:
TAS=RA—AS*PA+RA—ACS*PC−RSairB*PB eq. 5.24
To validate the equations 5.17, 5.21, and 5.24 shown above, the test values for RCA, RCB, RC, RB, RA, RSairB, RSairA, and RSairC, may be used to obtain values for RB_BS, RC_BCS, RC_CS, RB_CBS, RA_CAS, RA_AS, and RA_ACS, as shown below in tables 62 and 63.
Based on equations 5.17, 5.21, and 5.24, estimated values for TAS, TBS, and TCS may be obtained and compared to measured test results, as shown below in tables 64-69. Specifically, tables 64 and 65 show estimated and measured values for the test configuration shown in
Based on the data provided above, it can be seen that the estimated values for TA, TB, and TC are very close to the measured temperature values. Further, USL values and regression equations may be developed for the parameters RB_BS, RC_BCS, RB_CBS, RC_CS, RA_CAS, RA_AS, RA_CAS. As before, the USL values for these parameters can be used to avoid over-estimating these parameters and, thus, to avoid underestimating the junction temperatures.
The USL values for RCA, RA, RC, RBC, and RB are shown above in table 24. The USL values for RSairA, RSairB, and RSairC are shown above in tables 57-58. The USL values for RCA, RA, RC, RBC, RB, RSairA, RSairB, and RSairC can be used to determine USL values for RB_BS, RC_BCS, RB_CBS, RC_CS, RA_CAS, RA_AS, RA_CAS using equations 5.15, 5.16, 5.18, 5.19, 5.20, 5.22, and 5.23. For example, equation 5.15 can be used to obtain the USL values for RB_BS as shown below in table 71.
Equation 5.16 can be used to obtain the USL values for RC_BCS as shown below in table 72.
Equation 5.18 can be used to obtain the USL values for RB_CBS as shown below in table 73.
Equation 5.19 can be used to obtain the USL values for RC_CS as shown below in table 74.
Equation 5.20 can be used to obtain the USL values for RA_CAS as shown below in table 75.
Equation 5.22 can be used to obtain the USL values for RA_AS as shown below in table 76.
Equation 5.23 can be used to obtain the USL values for RA_ACS as shown below in table 77.
In an embodiment, regression techniques may be applied to the USL values obtained for the above parameters. Using the example data shown in tables 71 to 77 above, the following regression equations may be obtained.
RB—BS=0.0312+0.0693*EXP(−SCFM/24.88)+0.022*EXP(−SCFM/99.5) eq. 5.25
RC—BCS=−2.66E−2+0.5682*EXP(−SCFM/10.37)+0.0396*EXP(−SCFM/302) eq. 5.26
RB—CBS=−0.00929+0.31975*EXP(−SCFM/7.8) eq. 5.27
RC—CS=0.0299+0.0895*EXP(−SCFM/59.1)+0.087*EXP(−SCFM/13.5) eq. 5.28
RA—CAS=−2.19E−3−0.0418*EXP(−SCFM/18)−0.018*EXP(−SCFM/46.29) eq. 5.29
RA—AS=4.63E−02+0.1356*EXP(−SCFM/57)−0.0358*EXP(−SCFM/84.5) eq. 5.30
RA—ACS=−1.84E−2+0.0338*EXP(−SCFM/200.6)+0.5032*EXP(−SCFM/11.4) eq. 5.31
With regard to the thermal capacitances, for a thermal timing constant, τ, equal to 190 seconds at 150 SCFM and the test data, the data shown in table 78 can be provided.
It will be appreciated from table 78 that although −ve Rth shows cooling affect, −ve Cth does not have physical meaning, so these Cth's are zero, denoting an immediate affect (Cth=0 j/degr C.). Further, although the RA_CAS, using test data on 150 SCFM appears as a small number, its USL's for all SCFM are negative numbers. Thus, CA_CAS should also be taken as ZERO. This will make the capacitances of the interphases between phases equal to ZERO. The thermal impedance function derived herein may be used to determine real-time junction temperatures. For example, the thermal impedance functions described above may be programmed into the system controller 1702 (
The estimated junction temperatures may be used to control various aspects of the operation of the double H-bridge. In an embodiment, the applied load current may be modified based on the estimated junction temperatures, for example, by modifying the control signals used to drive the double H-bridge. In an embodiment, the estimated junction temperatures may be used in the process of controlling a traction motor, to which the double H-bridge is operably coupled for powering the motor. In an embodiment, the estimated junction temperatures may be used to control a cooling unit operably coupled to the double H-bridge. In an embodiment, the spatial, thermal, and/or electrical topology of the double H-bridge may be modified based on the estimated junction temperatures.
As shown in
In order to validate the results, TjB, TjC, TjA (denoted below as TjBS, TjCS and TjAS to indicate that results were obtained by estimating the sensor temperature) was estimated directly from Tair and compared to values obtained by estimating TSair and delta TBcase to Sensor, delta TCcase to Sensor and delta TAcase to Sensor. The results of the tests are shown below in table 79. For the test results shown in table 79, Vbatt=80 Volts and Tair=61 Degr C.
As can be seen from the data in table 79, the two sets of results are within a few degrees C., proving that the equations used to determine the junction temperatures provides a very good estimation of the thermal behavior of the double H-bridge converter. In an embodiment, the real-time, measured or estimated junction temperatures may be used by the double H-bridge controller to control the airflow rate of the double H-bridge's associated cooling unit.
The evolution of power electronic semiconductors has provided devices, such as IGBTs, with reduced power dissipation and increased junction temperature (Tj) capability. The latest generations of Isolating Gate Bipolar Transistors (IGBT's) have by far reduced power dissipation resulting in the ability to handle much more power. However, the improved power handling capabilities impose some additional constrains. As the upper temperature limit for operating the junction of the IGBT's increases, it also increases the thermal cycling of the device, which can result in reduced reliability long term over the long term absent additional safeguards.
Generally, there are two factors that restrict the thermal cycling capability of an IGBT, namely, the base plate soldering and the bond wires, both of which are subject to fatigue due to thermal cycling. The base plate soldering reliability depends, in part, on the material of the base plate. In an embodiment, the base plate soldering may use a metal matrix composite referred to as “AlSiC,” which includes an aluminum matrix with silicon carbide particles and provides more thermal cycling durability. To increase the durability of the aluminum wires interconnecting the chips inside the IGBT package, the wires may be coated.
As shown in
The system controller, based on the signals received by the ALC, estimates the required effective thermal resistances between the heatsink underneath each phase and the cooling air, RB* and RC*. The values of RB* and RC* may be similar to, but slightly larger than, the RBt and RCt values described above, because RBt and RCt were derived directly from the test data by allowing a three-sigma tolerance (Z=3). In order to be compatible with the rest of the simulation, RB* and RC* are derived from the USLs of RB, RBC, RC and RCA which had their standard deviation enlarged with the use of statistical modeling, resulting in larger values for RB* and RC*.
From equation 3.1:
TB−Tair=dTB=RB*PB+RBC*PC+RBA*PA
In equation 3.1, RBA equals zero, since there is no significant contribution of PA to dTB. Thus:
TB−Tair=dTB=RB*PB+RBC*PC
From equation 3.2:
TC−Tair=dTC=RC*PC+RBC*PB+RCA*PA
The USL values for RCA, RA, RC, RBC, and RB are shown in table 24. In order to simplify the computations for RB*, since RB>>RBC, the power Po=max(PB,PC) may be used for estimating the desired RthB_ha (desired RB*). Applying this simplification yields:
TB−Tair=RB**Po=RB*Po+RBC*Po=Po*(RB+RBC)
Solving for RB* yields:
RB*=RB+RBC eq. 7.1
Accordingly:
RB**Po=RB*Po+RBC*Po=Po*(RB+RBC)
Similarly, for RC*, since PA<max(PB, PC), RC* may be simplified to:
RC*=RC+RBC+RCA eq. 7.2
USL values for RB* and RC* may be developed, and are shown below in tables 81 and 82.
Based on the USL values from tables 81 and 82, it will be noticed that the USL values at the airflow rate of 0 SCFM appear to be out-liers. Regression equations describing the desired air flow rate as a function of RB* and RC* may be developed by applying regression techniques to the USL values for RC* and RB*. Applying such techniques to the exemplary USL data shown in tables 81 and 82 yields:
req.SCFM—B=36.43+769.62*EXP(−RB*/0.037) eq. 7.3
reqSCFM—C=34.95+591.2*EXP(−RC*/0.0465) eq. 7.4
In the above equations, SCFM_B, and SCFM_C are the airflow values desired for reliable operation of phase B and C, respectively. As shown in
The power dissipation of phase B or phase C may be referred to herein as PX, where X can equal B or C. The junction temperature of the phase A or phase B may be referred to herein as TjX, where X can equal A or B, and can be expressed as:
TjX=Tair+dTha+dTch+dTjc
In the above equation, dTha represents the temperature difference between the heat sink and the air, dTch represents the temperature difference between the IGBT case and the heat sink, and dTjc represents the temperature difference between the junction of the IGBT and its case. The parameters dTha and dTch can be expressed as follows:
dTha=PX*RX*
dTch=(PX/2)*0.018=PX*0.009
Thus, the equation for TjX can be expressed as:
TjX−Tair=PX*RX*+dTXjc+PX*0.009 eq. 7.5
Solving for RX* yields:
RX*=[(TjX−Tair)−dTXjc]/PX−0.009 eq. 7.6
Thus, the values of RB* and RC* can be computed based on the specified maximum thermal cycling guideline suitable for a particular application. In an embodiment, the max thermal cycling (TjX-Tair) in phase B may be specified to be approximately 64.5 degr C., and the max thermal cycling (TjX-Tair) in phase C may be specified to be approximately 68.5 degr C., which yields:
RB*=(64.5−dTBjc)/PB−0.009 eq. 7.7
RC*=(68.5−dTCjc)/PC−0.009 eq. 7.8
For explanation of the cycling levels used (64.5 and 68.5), see tables 84 and 85 below.
From the above, the worst-case steady state operating combination of Vlink, Ifield and Ibattery, can be determined as shown in tables 84 and 85 below. Specifically, the worst-case steady state operating combination for phase B is shown in table 84, and the worst-case steady state operating combination for phase C is shown in table 85.
In the example provided above, at any point of operation where the phase B dissipation is PB and the thermal difference between TBj and case B is dTBjc, the RB* value given by equation 7.7 will provide thermal cycling less than or equal to 64.5 degr C. Similarly, at any point of operation where the phase C dissipation is PC and the thermal difference between TCj and case C is dTCjc, the RC* value given by equation 7.7 will provide thermal cycling less than or equal to 68.5 degr C.
As shown in
To test the strategy described above, the system of
In table 86, the second column from the left indicates the available air flow from the equipment blower. As shown in table 86, if reqSCFM>available SCFM, then the available air flow is applied. It can also be seen from the data in table 86 that the desired airflow (reqSCFM) computed using equations 7.7 and 7.3 will be identical up to Vlink=1300V. However, using these two equations above 1300V it will result in over-estimating the desired airflow. However, above 1300V the blower will be operated close to, or at, the maximum airflow available, in other words, 198 SCFM. Based on these observations, the system shown in
To verify that the simplified system of
Based on the results shown in table 27, it can be seen that there is no difference in the “capability” of the double H-bridge using either system. However, the simplified system would have computed a greater required air flow for the two cases in which PB<PC (first two rows). However, since the max available air flow rate in this example is 198 SCFM, the two systems behaved identically.
Additional tests cases, shown below in table 88, were conducted to determine whether the system of
As shown in table 88, at some instances above 1300V, when PB<PC, the simplified system will overestimate the desired air flow rate, but at these high voltages the required air flow will generally be greater than the max available air flow rate of 198 SCFM. Examining different scenarios of Vlink=>1300V and PB<PC, the required air flow rates below 198 SCFM differed by less than 6-7 SCFM, which is insignificant. Additional tests were performed for the maximum (steady state) currents at a Vlink of 1500V, which are shown below in table 89.
As shown in table 89, since required air flow rate, reqSCFM, is at the upper specification limit of 198 SCFM, there is no change in the TjB, TjB-Tair, TjC, TjC-Tair between the two techniques. Based on the tests described above it can be seen that the simplified system of
Thermal Protection of IGBT of the Double H-Bridge
In embodiments, the system controller may be configured to thermally protect the IGBT's of the double H-bridge, in case of a system malfunction, such as failure of the blower providing the cooling air, air-leaks in the plenum, tunnel operation, and the like. For example, the load current may be de-rated as described below, to reduce thermal cycling.
As an example, under maximum steady state operating conditions in the double H-bridge, the maximum Tj-Tair may be specified as 68.5 degr C. This may occur, for example, at If=125 A, Ibatt=300 A, Tair=61 degr C. (Tamb=49 degr C.) at 1500 Vdc, TjC=129.41 degr C., and TChs=112.32 degr C. In this example, TChs is approximately 85% of TjC and it is measured by the temperature sensor 1700 (
Based on the exemplary values provided above, the system may be configured such that de-rating on Tj does not start until Tj>=137 degr C. When Tj-Tair is greater than 70 degr C. (calc Tj>131 degr C.), ALC (Auxiliary Logic Controller) may issue an indication that the IGBTs are getting hot, and no further action will be taken until Tj-Tair=76 degr C. (Tj=137 degr C.). In an embodiment, this stage will be omitted for countries with Tamb=55 degr C.
In an embodiment, the thermal cycling capability of the IGBTs is 75,000 thermal cycles of delta Tj=71 degr C. and 30,000 cycles of deltaTj=86 degr C. However, it will be appreciated that embodiments of the present techniques may include IGBTs with different thermal capabilities. Based on the deltaTj=86 degr C. and Tair=61 degr C., the double H-bridge controller may be configured to stop pulsing at Tj=147 degr C. or Tj-Tair=86 degr C. This provides a de-rating range shown below:
137 degr C.<=Tj<147 degr C., size 10 degr C., or
76 degr C.<=Tj-Tair<86 degr C., size 10 degr C.
In another example, for countries with Tamb=55 degr C. the double H-bridge controller may be configured to stop pulsing at 147 degr C. and the max delta cycling will be Tj-Tair=80 degr C. Notice that the absolute USL for Tj=150 degr C. This provides a de-rating range shown below:
137 degr C.<=Tj<147 degr C., size 10 degr C., or
70 degr C.<=Tj-Tair<80 degr C., size 10 degr C.
Embodiments of the present techniques may be better understood with reference to
It is to be understood that the above description is intended to be illustrative, and not restrictive. For example, the above-described embodiments (and/or aspects thereof) may be used in combination with each other. In addition, many modifications may be made to adapt a particular situation or material to the teachings of the invention without departing from its scope. While the dimensions and types of materials described herein are intended to illustrate embodiments of the invention, they are by no means limiting and are exemplary in nature. Other embodiments may be apparent upon reviewing the above description. The scope of the invention should, therefore, be determined with reference to the appended claims, along with the full scope of equivalents to which such claims are entitled.
In the appended claims, the terms “including” and “in which” are used as the plain-English equivalents of the respective terms “comprising” and “wherein.” Moreover, in the following claims, the terms “first,” “second,” “3rd,” “upper,” “lower,” “bottom,” “top,” “up,” “down,” etc. are used merely as labels, and are not intended to impose numerical or positional requirements on their objects. Further, the limitations of the following claims are not written in means-plus-function format and are not intended to be interpreted based on 35 U.S.C. §112, sixth paragraph, unless and until such claim limitations expressly use the phrase “means for” followed by a statement of function void of further structure.
As used herein, an element or step recited in the singular and proceeded with the word “a” or “an” should be understood as not excluding plural of said elements or steps, unless such exclusion is explicitly stated. Furthermore, references to “one embodiment” of the invention are not intended to be interpreted as excluding the existence of additional embodiments that also incorporate the recited features. Moreover, unless explicitly stated to the contrary, embodiments “comprising,” “including,” or “having” an element or a plurality of elements having a particular property may include additional such elements not having that property.
Since certain changes may be made in the above-described control method, without departing from the spirit and scope of the invention herein involved, it is intended that all of the subject matter of the above description or shown in the accompanying drawings shall be interpreted merely as examples illustrating the inventive concept herein and shall not be construed as limiting the invention.
| Number | Name | Date | Kind |
|---|---|---|---|
| 6160696 | Bailey et al. | Dec 2000 | A |
| 6601168 | Stancil et al. | Jul 2003 | B1 |
| 7334418 | Henry et al. | Feb 2008 | B2 |
| 7826216 | Moss | Nov 2010 | B2 |
| 20060119177 | Kumar et al. | Jun 2006 | A1 |
| 20080238195 | Shaver et al. | Oct 2008 | A1 |
| 20090072770 | Son et al. | Mar 2009 | A1 |
| 20090315328 | Kumar | Dec 2009 | A1 |
| Number | Date | Country |
|---|---|---|
| 1826897 | Aug 2007 | EP |
| 2007215250 | Aug 2007 | JP |
| Entry |
|---|
| Search Report and Written Opinion from corresponding PCT Application No. PCT/US2012/025463 dated Aug. 29, 2012. |
| Search Report and Written Opinion from corresponding PCT Application No. PCT/US2012/025452 dated Apr. 10, 2013. |
| Number | Date | Country | |
|---|---|---|---|
| 20120218710 A1 | Aug 2012 | US |
