Installation Considerations for Multi - Motor AC Drives & Filters
Used In Metal Industry Applications
G. Skibinski*, D. Dahl*, K. Pierce*, R. Freed**, and D. Gilbert**
*Rockwell Automation
6400 W. Enterprise Drive
Mequon, WI 53092
(414)–512–7151 (414)–512–8300 fax
**Bricmont Inc.
395 Valley Brook Road
McMurray, Pa 15317-3397
(412)-941-3150 Ext. 236
Abstract- This paper investigates the zero sequence current that
develops in the ground circuit of a multi-motor drive
application. The motivation was to reduce the magnitude of
common mode electrical noise in the system ground grid and to
reduce occurrence of ground fault sensor trips as a result of this
noise. The phenomenon of line to ground cable charging current
(Ilg) during drive switching is a main component of zero
sequence current and is discussed with basic equations
describing its magnitude. Results of a high frequency
characterization study of all components in the zero sequence
path was required to obtain estimates of surge impedance for
these equations. Measured surge impedance was also used in a
simulation program designed to predict Ilg magnitude for
various system conditions. A PWM output filter and an input
isolation transformer with the neutral high resistance grounded
were two solutions investigated to reduce the zero sequence
current that were both simulated and measured on site.
Gate Bipolar Transistor (IGBT) inverters with 50 ns switch times
and higher switching rates (2 kHz to 12 kHz) requires a careful
investigation of the zero sequence cable charging current problem
for multi-motor drives using long drive output cables. The IGBT is a
gate voltage driven device which can source high magnitudes of
transient output current, limited only by external drive impedance.
The fast risetimes create a line to ground impulse voltage source
during switching, which tends to maximize the peak line to ground
current seen. It is shown that each long cable leaving the drive may
have a 3 Apk line to line current as well as an 8 Apk current from
phase wire capacitance to ground. A plant system ground current
problem now exists with the newer drives because all 163 motors
are switched synchronously resulting in 8 Apk times 163 cables or
1,300 amps peak transient current to ground possible during every
switch instant! This current is associated with approximately 8 miles
of three phase drive output cable.
This paper describes the use of a special PWM output filter to
reduce the effect zero sequence current on drive ground fault trips,
system ground fault indicator mis-operation and increased Electro
Magnetic Interference (EMI) noise in the plant ground. The
installation of the special PWM output filter substantially reduced
the peak ground current for synchronous operation of 72 one hp
motors. However, the drive ground fault trip returned as power to
the last 91 one hp motors was applied. To circumvent this problem,
the original plant one line power diagram was changed from a single
2.5 MVA solid ground transformer feeding three 250 hp drives, to a
solid ground 2.5 MVA plant transformer feeding separate 250 kVA
resistive ground isolation transformers for each 250 hp drive as
shown in Fig. 1. The remaining sections discuss the cable charge
phenomenon, equations describing its magnitude, characterization
results of component high frequency surge impedance’s used in
determining charge current magnitude and system simulation.
Measured field site data is compared to system simulation results.
Solutions to the cable charge phenomenon and pre-installation
application guidelines are provided to insure successful utilization of
the multi-motor drive topology that is typically used in the metals
industry.
I. INTRODUCTION
Over the years, the metals industry has found that a single
Variable Frequency ac Drive (VFD) synchronously powering
multiple ac induction motors is both an economical and flexible
configuration for steel processing requirements. In a typical
application, a single high horsepower ac drive may synchronously
power up to 163 low hp induction motors on the conveyer of a 400
foot long steel process tunnel furnace, as shown in Fig. 1 and Fig. 2.
The use of drive output contactors on each motor allows flexibility
in selecting the number of motors required for a given iron slab run
through the 400 ft long furnace. In normal operation, each 250 hp
drive may use 30 to 80 motors to control a slab in a particular
section of the oven. The slab is transferred to the next drive section
and caught on the fly where its particular speed and temperature
may be readjusted for steel processing requirements. The output
contactor switching arrangement also provides process uptime and
redundancy in the event of drive, cable or motor failure. A failed
cable or motor may be quickly switched out, while a failed drive
may have its load motors transferred to another drive. Finally, a
single 250 hp drive is required to power all 163 motors if a large
slab is processed.
The inherent length of the tunnel furnace dictates long drive to
motor output cables with substantial line to ground capacitance as
shown in Fig. 3. This creates the potential for large zero sequence
cable charging currents during switching instants of the VFD output
waveform. Previous generation six step drive inverters used Bipolar
Junction Transistor (BJT) semiconductors with slow 1-2 µs switch
times which tended to reduce the peak line to ground current. Also,
BJTs were gate current driven devices which beta limited the
maximum output current allowed. The BJT would stay in the active
region under beta limit control during switching, until either the
cable was fully charged or a transistor desaturation protection circuit
was faulted. Some BJT applications upsized the drive hp
requirements or would require a simple 3 phase output line reactor
to prevent faults induced by cable charge currents. Use of Insulated
II. CABLE CHARGING CURRENT PHENOMENON
Cable charge current consists of transient line to line (Ill) and line
to ground (Ilg) components. Fig. 4 defines these current paths in
motor & Power Equipment (PE) ground wires for a grounded
conduit configuration. Wires are represented by distributed line
inductance (Lo1), distributed line to line stray capacitance (Cll) and
distributed line to conduit ground cable capacitance Clg. The Clg
component additionally consists of stator winding to motor PE
frame ground capacitance (Csg).
A. Line to Line Cable Charge Current
Each IGBT switching interval sets a dv/dt transition of the line to
line PWM voltage pulse (Vll) in Fig. 5 and induces a transient line to
line cable charge current Ill from the drive dc bus capacitor through
0-7803-4943-1/98/$10.00 (c) 1998 IEEE
the gated IGBT. The traveling Ill wave enters distributed cable Lol ,
into Cll and returns through another phase directly back to the drive
bus capacitor. The Vll zero voltage dwell time prior to the next dv/dt
transition to the dc bus voltage (Vbus) level dictates whether cable Cll
is fully discharged to zero volts following the initial transient.
Initially uncharged cable conditions are typical for drives with low
PWM carrier frequencies (fc) or short output cables. High fc or long
cable conditions may have a positive or negative residual cable
voltage (Vres) prior to Vbus transitions, depending on interaction of
the PWM modulator and output cable characteristics [1]. Peak
transient Ill cable charge current associated with Vll traveling wave
voltage is estimated in (1) using cable line to line surge impedance
(Zll), Vbus and a (+/-Vres), dependent on whether Vres aids or opposes
the incoming Vbus value. Section III further discusses typical
magnitudes for Zll.
I ll =
V bus ( + / − ) V
res
Z ll
where Z
ll =
Lo1
Cll
(1)
B. Line to Ground Cable Charge Current
Each IGBT switching transition in a given phase also sources a
line to ground cable charging current path in Fig. 4 from line
inductance Lo1 through Clg and Csg. The zero sequence voltage
waveform (Vng) at the drive PWM output is the source of Ilg current.
Fig. 6 shows Vng and Ilg waveforms for a single 480V drive (Vbus =
650 Vdc) operating at fc = 4 kHz and output 30 Hz frequency and
connected to a single motor cable of 300 ft. Vng was measured
between the neutral point of a wye configured 1 Meg-ohm resistor
network connected to phase A, B, C and drive PE ground point. Vng
contains a dc bus related 180 Hz ripple voltage and 4 kHz
modulation voltage component. Fig. 7 is an expanded Vng version
showing the 4 kHz step like modulation component along with
associated Ilg transient current at each IGBT transition. Ilg is the
common mode line to ground current summation of all 3 phases,
measured by passing all three motor phase wires thru a Pearson high
frequency current transformer. In contrast to the Vll waveform of
Fig. 5, there is no dwell time in Fig. 7 Vng waveform. Highest Ilg
transients (8 Apk) in Fig. 7 occur when capacitance Clg and Csg have
previously attained a steady state charge of opposite polarity to the
incoming step magnitude change. Assuming the theoretical neutral
of the PWM inverter is the Vbus midpoint, then the maximum line to
ground step magnitude change is (Vbus /2) at the drive output. Fig. 7
shows highest Ilg transients occur when Clg is charged to a -250V
(-Vbus /3) plateau and a charge reversal step of Vbus /2 = 325 V is
applied. Because of the charge reversal phenomenon, the maximum
transient line to ground forcing voltage is roughly Vbus. The Ilg
magnitude is limited by line to ground surge impedance (Zlg) in (2).
Section III further discusses typical magnitudes for Zlg.
I lg
=
~
V bus
where
Z lg
Z lg
=
Lo1
Clg
(2)
C. Zero Sequence Current Paths
Fig. 8 shows system Ilg current paths taken for a configuration
consisting of 3 phase output wires plus ground wire enclosed in a
conduit and an input tray cable from a wye grounded feeder
transformer. The conduit is bonded to drive cabinet and motor
junction box and the green ground PE wire is connected to ground
stud in the motor junction box and drive cabinet PE bus. Transient
Ilg current sourced from the drive flows thru cable capacitance to the
grounded conduit wall, to green PE ground wire and partly thru
motor stator winding capacitance to frame ground. The conduit, and
to a lesser extent the internal green PE ground wire, absorb most Ilg
current and return it back to the drive out of the ground grid, thereby
reducing "ground noise" for the length of Potential #1 - Potential #2
run. However, a conduit may have accidental contact with grid
ground structure due to straps, support, etc. that bypasses Ilg back
into the ground grid. How Ilg current divides between the conduit,
green wire or ground grid is dependent on the variable and
unpredictable ac resistance characteristics of earth ground at the
application site. Upon arriving at drive PE, Ilg must enter the ground
grid, since a physical connection between drive PE ground and
power structure does not exist. Ilg must bypass drive PE ground, and
remain in the ground grid until the feeder transformer secondary
grounded neutral (Xo) is found. At this point a metallic path back to
the drive source can occur on input phase R, S or T. Once inside the
drive, Ilg current path selects the bridge rectifier diode that is
conducting back to the dc bus source.
The Ilg ground current path spans both output and input cable
lengths, in contrast to the Ill current path which was solely confined
to the output wires. It is also evident that placing the feeder
transformer close to the drive results in less ground grid noise.
Using conduit with a PE ground wire bonded from feeder
transformer Xo to the drive input is also beneficial in reducing
ground grid noise. It is seen in Fig. 8 that one method to reduce Ilg is
to insert a high resistance type grounding resistor in the series Ilg
path of the transformer Xo lead. This is discussed in Section IV.
D. Zero Sequence Current Problems
Fig. 9 shows Ilg and Ill current spikes adding to the fundamental
output phase current of a single low hp motor. In this multi-motor
application, Ilg and Ill currents travel synchronously down each of the
160 motor cables, so that peak zero sequence current magnitude
summed in the ground grid becomes enormous. The drive Ground
Fault Circuit prevented continuous operation because of the sum
total of the Ilg magnitude. An additional problem encountered is that
the RMS value of the Ilg and Ill currents must be added to the one hp
rated load current value to prevent the output overload heater from
malfunctioning. The RMS value of the Ilg and Ill currents may be
decreased by reducing fc. However, peak Ilg current remains
unchanged and may still cause system noise problems. Common
mode noise in the system ground grid created by the zero sequence
current may also effect sensitive electronic equipment referenced to
ground and is discussed in [2-5]. The solution to these problems was
the addition of a special PWM output filter and an input high
resistance ground isolation transformer for each drive discussed in
Section V.
III.
CHARACTERIZATION OF SYSTEM COMPONENTS
This section characterizes the high frequency surge impedance of
components that determine peak current magnitudes for Ill and Ilg in
(1) and (2). The high frequency impedance of other system
components in the Ilg path are also characterized for use in Section
IV simulation. Simulation is used for analysis as well as design of
the component parameters required to solve the cable charge current
problem. Component impedance was measured with HP 4284A and
HP 4285A R-L-C Impedance Analyzers with frequency ranges of 20
Hz - 1 MHz and 75 kHz to 30 MHz, respectively. Component surge
impedance at the specific frequencies corresponding to drive
transition time (fn) and Ilg oscillation frequency (fo) are of
importance. Equivalent frequency fn is defined in [3] and Fig. 5 as
fn = 1/π trise and is ~ 1.3 MHz for the 250 hp output voltage
risetime transition of 250 ns. The Zll at fn is used to calculate Ill and
Ilg in (1) and (2). The Ilg oscillation frequency fo typically varies
between 100 kHz and 1 MHz.
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A.
Normal Mode Zll Impedance of Wires in Conduit
Normal mode and common mode impedance of the application’s
#12 AWG Poly Vinyl Chloride (PVC) wires in a grounded conduit
was experimentally measured with impedance analyzers for a 10 ft
conduit section. Measured surge impedance Zll between two
conductors was determined by measuring open circuit capacitance
Coc between two phase wires and by measuring short circuit
inductance Lsc between the two wires at the conduit input side with
the same two wires shorted at the output side of the conduit [6]. The
60 Hz capacitance Coc between phases was 40 pF and is much lower
than that of a tightly bundled cable since the inherent air space
between wire insulation reduces the “effective dielectric” constant”
closer to that of air (εr = 1) rather than the PVC insulation (εr = 5.6)
value. It was important to ground the test conduit during normal
mode measurements. Otherwise, the Coc value doubled to an
erroneous 80 pF. value. The 60 Hz Coc value variation with
frequency in Fig. 10 decreases ~ 10 % at 100 kHz and ~ 50 %
during the drive transition equivalent frequency of 1.3 MHz.
L sc
(3)
Coc
High frequency inductance and resistance variation in Fig. 10
follows the classical skin and proximity effect profile for two wires
with equal and opposite current flow that are semi-adjacent to each
other. The inductance decreases to 60 % to 70 % of the 60 Hz value
over the 100 kHz to 2 MHz range. The [ac/dc] resistance ratio is
[6x] or 0.24 Ω at 100 kHz and [40x] or 1.6 Ω at 2 MHz.
Fig. 12 shows line to line surge impedance Zll variation with
frequency calculated from (3) and measured Lsc - Coc data of Fig. 10.
The Zll = 300 Ω at fn , so peak Ill transient using (1) is ~ 2 Apk for a
650 Vdc bus. While the 1.6 Ω ac resistance at fn provides transient
damping, it is seen the 300 Ω surge impedance approximation of a
lossless line dictates Ill magnitude. Peak line to line charge current
magnitude of PVC wires in a conduit is 4x to 5x lower than for
tightly bundled armor cable, since Zll range for these cables is ~
50 Ω - 80 Ω [6]. The 250 hp drive without any output filter must
transiently source an Ill = 2.2 Apk*163 cables = 358 Apk.
Z ll
B.
=
Common Mode Zlg Impedance of Wires in Conduit
Measured surge impedance Zlg between phase conductor and
conduit ground was determined by measuring open circuit
capacitance Coc between phase wire and the conduit and by
measuring short circuit inductance Lsc between phase A wire input
and conduit input with opposite end of phase A bonded to the
conduit output side. The test conduit was floating during common
mode measurements. The 60 Hz capacitance Coc between phase and
conduit was 250 pF and is much higher than the line to line value.
The 60 Hz Coc value variation with frequency in Fig. 11 decreases ~
15 % at 100 kHz and also ~ 15 % during the drive transition
equivalent frequency of 1.3 MHz.
High frequency inductance and resistance variation in Fig. 11
follows the classical skin and proximity effect profile for a common
mode Ilg current path that goes down a copper phase wire and
returns in the opposite direction down a coaxial steel tube.
Inductance decreases to 40% of the 60 Hz value over the 100 kHz to
2 MHz range. The [ac/dc] resistance ratio is [4x] or 0.3 Ω at 100
kHz and [20x] or 1.5 Ω at 2 MHz.
Fig. 12 shows line to line surge impedance Zlg variation with
frequency calculated using (3) but measured Lsc - Coc data of Fig. 11.
The Zlg = 80 Ω at fn , so peak Ilg transient using (2) is ~ 8 Apk for a
650 Vdc bus. At frequencies > 5 MHz, the Zlg of Fig. 12 tends toward
a 50 Ω coaxial value. The 1.5 Ω ac resistance at fn for the measured
10 ft section provides transient damping and is later seen to affect Ilg
magnitude. The 250 hp drive without any output filter or damping
considered must theoretically transiently source an Ilg = 8 Apk*163
cables = 1,300 Apk.
Common mode surge impedance testing with the insulated PE
ground wire bonded to both ends of the conduit was also performed.
Results show that Zlg of Fig. 12 decreased by only 5% over the 100
kHz to 5 MHz range. This implies that the PE wire does not carry
the high frequency zero sequence current. Fig. 13 shows test results
of an isolated 300 ft section which plots Ilg components in the
conduit and the PE ground wire during a Fig. 7 step transition in
common mode voltage. Fig. 13 verifies that the conduit, although
made of steel, looks like a low inductance coaxial tube to high
frequency, while the insulated PE wire appears a high value inductor
at high frequency.
The transient Ilg component of 1,300 Apk during switching will be
modified to lower value dependent on the surge impedance of the
input transformer, output filter and motor components in the Ilg
ground path, which spans both output cable and input cable lengths.
The Ill = 358 Apk component estimate will also be modified to
lower value dependent only on the output filter surge impedance
addition.
Input Transformer Impedance
C.
The equivalent circuit R-L-C parameters at high frequency are
estimated in Table 1 for both 2.5 MVA and 250 kVA transformers
used. The R(60Hz) - L(60Hz) parameters of Table 1 are per phase values
referred to the secondary side and determined using Fig. 14
transformer (X/R) ratio and (%Z) data estimates @ 60 Hz and
equations (4) through (6). The L(60Hz) parameter corresponds to air
core coil leakage reactance of both primary and secondary coils.
V base
=
V ll ( ac )
3
;I
base
Z phase = % Z * Z base =
X
R
=
K from Fig.14
=
3 phaseVA rating
3 V
V base
=
; Z
base
ll ( ac )
2
2
Xl + R =
(4)
I base

2  2
 K + 1 R


( )
;L
phase
=
(5)
Xl
2π f
(6)
Determining accurate estimates for high frequency R-L parameters
(R(1MHz) - L(1MHz)) at fn would itself be study subject outside the scope
of this paper. Also, time did not permit on-site testing of the two
transformers used. High frequency R-L approximations at fn were
determined by adjusting R(60Hz) - L(60Hz) with the two wire results of
Fig. 10 with (Rac/Rdc = 40x) pu factor and (Lac/Ldc = 0.7x) pu
factor. It is assumed ½ of the leakage is in the primary and ½ in the
secondary for both transformers. Fig. 8 shows that for the 2.5 MVA
transformer, only the leakage inductance and resistance of the
secondary side conduct zero sequence current Ilg, and that Ilg, may
flow on all three phases in parallel ( 1/3 pu phase value) back to the
drive. Thus, the 2.5 MVA (R(1MHz) - L(1MHz) ) parameters are
multiplied by (1/6) to get (R1(Equiv) - L1(Equiv)) parameters used in the
system P-SPICE model of Fig. 18 for Ilg magnitude determination.
Table 1
Equivalent R-L-C Parameters for 250 kVA & 2.5 MVA transformers
________________________________________________________
kVA R(60Hz) L(60Hz) R(1MHz) L(1MHz)
R(Eqiuv) L(Equiv) C(Equiv)
________________________________________________________
10 µΗ
250 7.68 mΩ 86 µΗ 307 mΩ 60 µΗ 50 mΩ
20 nF
R2 & R4 L2 & L3
2 µΗ
2,500 755 µΩ 14 µΗ 30 mΩ 10 µΗ
5 mΩ
L1
R1
________________________________________________________
In a similar approach, L2(Equiv) and L3(Equiv), each represents the zero
sequence high frequency primary and secondary non-saturable air
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core leakage inductance of the 250 kVA 1:1 isolation transformer
and R2(Equiv) , R4(Equiv), each represents the zero sequence high
frequency primary and secondary coil resistance. The C(Equiv)
parameter was measured between a primary phase lead and Xo of
the secondary side on the 250 kVA transformer at the field site. The
C(Equiv) parameter simulates the inherent high frequency coupling
capacitor between primary and secondary coils that can partially
transfer part of the Ilg current.
D.
Output Filter Impedance
The primary filter component of Fig. 15 consists of a 250 hp 3
phase reactor rated at 5% impedance. The reactor is beneficial in
reducing both Ill and Ilg charge current components. The apparent 60
Hz inductance at rated current is determined using (7) and by
measuring phase resistance Rphase and applying Vrms until rated Irms
occurs.
V rms
I rms
= Z phase =
X phase
2
2
X phase + R phase ;
=
L phase
2π f
IV. SIMULATION OF FIELD CONDITIONS
(7)
The 5% Z rating is the highest inductance allowed that does not
significantly decrease the 60 Hz output voltage to the motor.
However, the apparent 60 Hz inductance of these reactors drops
significantly at the important fn frequency because of the magnetic
skin effect of the core laminations. Fig. 16 shows impedance
analyzer results indicating only 18% of the 60 Hz inductance (~
20 µΗ) exists in the 1 MHz to 2 MHz region where it is desirable to
maintain inductance to limit Ilg during drive switching. Reactor core
loss resistor Rcore was measured at fn to be 250 Ω. The ac skin effect
resistance of each phase coil at fn was measured to be 60 mΩ. An
external resistor Rdamp was added to each phase to provide additional
damping in the output circuit. Only phase values are used in the
output circuit simulation since Ilg is sourced from one switching
instant in an inverter phase.
A common mode core made of high frequency ferrite material in
Fig. 15 was also added in series to help compensate for the drop in
iron core inductance during switching.
E.
Motor Common Mode Surge Impedance
The transient Vll waveform entering the motor is propagated and
attenuated through the stator line to line winding, which is modeled
as a distributed transmission line [4]. There have been simple and
more complex methods attempts to model the high frequency line to
line surge impedance characteristics of the motor winding [7,8].
However, there is little literature on the common mode surge
impedance of the motor. Of interest in calculating Ilg magnitude is
the motor common mode surge impedance when the zero sequence
traveling wave enters the motor winding to frame ground. The Zlg of
the motor winding was measured using the designated R-L-C
analyzers and appears as a phase to ground capacitor with
impedance Xc = (1 / 2πf Cmg) at 60 Hz. Fig. 17 shows the measured
60 Hz Cmg values for various motor hp sizes. The Cmg of a 1 hp
motor is between 1 nF to 3 nF. The Zlg at the fn frequency is also
highly capacitive, but the Cmg value is reduced and the skin effect
resistance of the winding becomes evident. The P-SPICE model
used at fn, , C4 in Fig. 18, is thus a simple capacitor to ground with a
series resistor accounting for ac skin effect resistance of the
winding.
F.
µΗ * 40% effect skin factor per 10 ft length). However, all lines are
in parallel for 161 motors so the net lumped cable inductance is 0.58
µΗ. The values indicate the high frequency inductance of the filter
dominates over any other component.
The total lumped capacitance in the series Ilg path at fn is ~ Cmg *
161 motors = 484 nF plus cable capacitance. The output cable
capacitance is estimated as a lump sum of 10.6 nF per motor line for
500 ft length (Fig. 11 – 250 pF * 85% effect skin factor per 10 ft
length). However, all lines are in parallel for 161 motors so the net
lumped cable capacitance is 1.71 µF. The values indicate cable
capacitance dominates over motor to ground capacitance.
An Ilg estimate using a lumped parameter surge impedance
model of the components in the Ilg path is calculated using a lumped
Lol of 24.58 µΗ and 2.19 µF for lumped capacitance Clg in (1). This
results in Zlg of 3.35 ohms and peak Ilg of 190 Apk without taking
into the effect of damping resistance.
Ilg Estimate using System Surge Impedance
The total inductance in the series Ilg path at fn is ~ 20 µΗ for the
filter inductor, 2 µΗ for the 2.5 MVA transformer, 2 µΗ stray
inductance in the drive. The output cable inductance is estimated as
a lump sum of 95 µΗ per motor line for 500 ft length (Fig. 11; 4.75
A.
Zero Sequence Equivalent Circuit
Fig. 18 shows a P-SPICE equivalent circuit model of all
components in the Ilg zero sequence path defined at fn. The 2.5 MVA
solid Xo ground transformer parameters R1 & L1 were defined
previously. The 250 kVA transformer parameters R2, R4 and L2, L3
were defined previously and resistor R3 = 250 Ω that is connected to
Xo. The 250 kVA transformer and R3 combination was a final
solution that was not initially at the application site. The input
conduit consisted of 75 ft of 320 amp conductors in a large conduit
with lossless transmission line parameters L4 and C2. Resistor R5
simulates high frequency skin effect resistance losses in the cable.
The drive is represented by a step voltage change through an IGBT
with ON state resistance and some stray inductance in the drive bus
work. The output filter lumped parameters R6 & L5 were defined
previously. A P-SPICE lossless transmission line model for the
single output cable was previously defined using C3 & L6 values per
meter. Resistor R7 is a lumped value for the cable skin effect loss. A
more sophisticated transmission line model that models distributed
cable losses is detailed in [9] using the SIMULINK program .
Capacitor C4 is the line to ground Cmg of 3 nF for one motor. The
remaining 162 motors are grouped into one lossless transmission
line, loss resistor and motor model. C6 is the line to ground Cmg of
162 motors or 480 nF. The skin effect resistor R8 is a lumped value
calculated as (R7 /162) and line inductance is (L6 /162).
The value of the model lies in its ability to predict current for any
system combination of components without expensive field testing.
Fig. 19 through Fig. 22 simulation results are for various cases of
synchronously switching 163 motors at once and measuring Ilg.
B.
Charging Current Simulation
Fig. 19a simulation results are shown with the 2.5 MVA and 250
kVA transformer, input conduits, output filter and cable skin effect
resistor R7 and R8 all removed from the circuit. An Ilg value of 1,310
Apk with 500 kHz cable oscillation ring frequency that never decays
is shown and agrees with the Section III-B 1,300 Apk prediction.
Fig. 19b shows just the addition of cable ac skin effect resistors R7
and R8 alone was enough to decrease peak Ilg magnitude to 489 Apk
at the same 500 kHz oscillation frequency of the cable parameters,
but now with a decay to zero in 6 µs.
Fig. 19c shows the addition of the 2.5 MVA transformer, input
conduit and non-ideal IGBT drive model, to the existing ac cable
skin effect resistors R7 and R8. These additions were enough to
decrease peak Ilg magnitude to 319 Apk, at much different and lower
42 kHz oscillation frequency, but now with a much longer decay to
zero in over 100 µs.
0-7803-4943-1/98/$10.00 (c) 1998 IEEE
C.
Simulation with Output Filter Added
Fig. 20 shows the addition of the 2.5 MVA transformer, input
conduit, non ideal IGBT drive model and PWM output filter models
along with existing ac cable skin effect resistor R7 and R8 . This
condition simulates the system originally shipped to the customer
and corresponds to the brute force peak Ilg 190 Apk estimate of
Section III-F. The peak Ilg magnitude simulated is decreased to 110
Apk at a lower 14 kHz oscillation frequency, but now with a much
longer decay to zero in over 250 µs as compared to Fig. 19c. The
graph of Fig. 20 also shows a reduced Ilg component that goes
through the motor stator winding to ground. The output line
common mode surge impedance is predominant in determining peak
Ilg out of the drive. The addition of filter inductance may reduce Ilg,
but it also lengthens the oscillation period, which may lead to higher
peak Ilg, if it has not decayed to zero before the next drive switching.
The drive ground fault sensor trips at exactly 100 amp and also
caused random trips due to the 110 Apk.
D. Simulation with Isolation Transformer & Output Filter
Fig. 21 shows the simulation of the final solution involving all of
Fig. 18. The addition of a standard high resistance grounding
resistor R3 = 250 Ω connected to Xo of the 250 kVA one to one
isolation transformer added provided the substantial damping
required in the zero sequence ground path. The leakage reactance
provided extra non-saturable inductance’s to limit current. The
primary to secondary series capacitance actually changed the circuit
oscillation frequency. Fig. 20 shows the 250 kVA transformer
primary zero sequence current still conducts a large percentage of
the Ilg component at the drive output. The equivalent circuit of Fig.
18 shows that R3 is actually damping resistor in parallel with the 2.5
MVA secondary side and the 250 kVA primary side. The peak Ilg
magnitude simulated is decreased to a tolerable 24 Apk at a higher
120 kHz oscillation frequency, and now with a drastically reduced
decay to zero in 20 µs to 40 µs as compared to Fig. 20.
E. Simulation with Grounded Transformer & Output Filter
The effect of varying R3 = 0 Ω to represent a grounded wye 250
kVA isolation transformer with all the components of Fig. 18 intact
was investigated by simulation. Fig. 22 shows the peak Ilg
magnitude in the 2.5 MVA secondary side and the 250 kVA primary
side goes to zero. The Ilg oscillation frequency reverts back to 18
kHz, since C1 is essentially shorted. The peak magnitude increases
to an unacceptable 141 Apk with a 250 µs decay time due to no
damping resistance in the zero sequence path. Thus, a solid
grounded isolation transformer is unacceptable for drive ground
fault indicator operation.
V.
oscillation frequency and 20 µs decay time agree well with the
simulated 24 Apk, 120 kHz oscillation frequency and 20 µs to 30 µs
decay time.
Fig. 24 is a long term time scale of Fig. 23a showing both Ilg
common mode current and phase current at the drive output that was
measured at the field site for 163 motors starting up to 30 Hz output
frequency. Note the high RMS content of the zero sequence current.
The phase current does not represent the Ilg spikes due to aliasing at
5 ms per division.
Fig. 25 is a long term time scale of Fig. 23b showing the PE
insulated green ground wire Ilg common mode current component
vs. the Vll switching waveform that was measured at the field site for
163 motors at 30 Hz output frequency. Note the lower RMS content
of the PE zero sequence current as compared to the total common
mode current of 32 Apk. This agrees with previous lab
measurements showing that most of Ilg current conducts within the
conduit.
VI. CONCLUSION
This paper described how line to ground cable charging current is
generated, the paths it takes and the system impact when there are
multiple parallel output cables attached to single drive output. High
frequency measurement techniques were outlined on how to
determine the impedance’s of the components in the path of the
cable and motor line to ground charging current. Knowing these
impedance’s, simple equations were shown to give rough estimates
of the cable charge current magnitude. A P-SPICE model of the zero
sequence circuit was shown to be accurate tool and had good
agreement with measured data taken at the field site.
In general, the use of an output filter alone, consisting of mostly
inductance, tends to reduce the cable charge magnitude but spreads
the resulting oscillation over a longer time. Selection of a filter
damping resistor, in parallel with the output filter, would not
significantly affect the oscillation decay time.
The use of a high resistance grounded input isolation transformer
in addition to the output filter, added extra line inductance and
tended to reduce the cable charge magnitude. The transformer high
resistance ground resistor in series with the ground path through Xo
provided a place for a filter damping resistor, that is not in the main
normal mode current path, and which could significantly decrease
the oscillation decay time and peak charge current to ground
magnitude.
The use of a solidly grounded input isolation transformer, with or
without an output filter, does not significantly decrease the
oscillation decay time nor the peak charge current to ground
magnitude.
ACKNOWLEDGMENT
The authors wish to thank D. Schlegel, J. Pankau and H. Jelinik for
impedance measurement and paper support activities.
FIELD SITE MEASUREMENTS
[1]
Fig. 23a shows Ilg common mode current at the drive output that
was measured at the field site for 163 motors starting up to 30 Hz
output frequency. Fig. 23a corresponds to conditions of a 2.5 MVA
source transformer with a PWM output filter that was simulated in
Section IV-C and Fig. 20. The 120 Apk, 14 kHz oscillation
frequency and 250 µs decay time agree well with the simulated 110
Apk, 15 kHz oscillation frequency and 250 µs decay time.
Fig. 23b also shows measured Ilg common mode current at the
drive output with six slabs of steel on the rollers. The Ilg was
measured at the field site for 163 motors starting and running to an
output frequency of 12 Hz. Fig. 23b corresponds to conditions of a
2.5 MVA source transformer, 250 kVa isolation transformer with
250 Ω wye ground resistor and the PWM output filter that was
simulated in Section IV-D and Fig. 21. The 32 Apk, 120 kHz
[2]
[3]
[4]
[5]
[6]
[7]
[8]
[9]
REFERENCES
Kerkman,R.,Leggate,D.,Skibinski,G.,"Interaction of drive modulation
& cable parameters on AC motor transients",IEEE-IAS-1997 IAS Conf
G. Skibinski, J. Pankau, & W. Maslowski,"Installation considerations
for IGBT ac drives",1997 IEEE Annual Textile,Fiber & Film Ind. Conf
G. Skibinski, J. Pankau, R. Sladky, J. Campbell, “Generation, control
and regulation of EMI from ac drives, IEEE-IAS-1997 Annual Conf .
Anderson,Kerkman,Saunders,Schlegel, and Skibinski, "Modern Drives
Application Issues & Solutions Tutorial", 1996 IEEE-IAS-(PCIC),PA
Rendusara,D.,Enjeti,P.,“Inverter output filter reduces common mode &
differential mode dv/dt at motor terminals in PWM drives, 1998 PESC
Skibinski, G., "Design methodology of a cable terminator to reduce
reflected voltage on AC motors", IEEE Ind. Appl. Soc. Conf., 1996
Skibinski,Kerkman,Leggate,Pankau,Schlegel,“Reflected wave modeling
techniques for PWM ac motor drives,”1998 IEEE-APEC, Anaheim,CA
Grandi,G.,Casade,D., Reggiani,U.,” Analysis of common mode &
differential mode hf current in PWM fed ac motors”,1998 PESC, Japan
Leggate, Pankau, Schlegel, Kerkman, Skibinski, ”Reflected waves and
their associated current, 1998 IEEE-IAS annual Conf.
0-7803-4943-1/98/$10.00 (c) 1998 IEEE
Fig. 2. Picture showing tunnel furnace section with motors
Fig. 3. Picture showing potential for high zero sequence currents
flowing in the 163 multiple conduits and wireway
Vng
Fig. 1. One line diagram of steel process furnace (400 ft long) showing a 2.5
MVA feeder transformer, three 250 Hp IGBT PWM drives, output filters &
isolation transformers, output contactors, wireways, and conduits to 163 one
hp motors.
IM O T O R
A
L 01
B
Drive
C
L 02
PE
L 03
C ll
C lg
Motor
C lg
C mg
Fig. 6. Measured Vng & Ilg of a 480V drive @ 30 Hz, 300' of motor cable
[ "0" at offest 4 div; 2 ms/div; 2 A/div; 100V/div]
Conduit
Fig. 4. Line to line Ill and line to ground Ilg cable charging current paths
V ll
Ilg
fc
t rise
tfall
f n = .3 18 / t rise
DC Bus
τ
t
I ll
t
Vng
Fig. 5. Line to line Vll PWM voltage and Ill charging current
Fig. 7. Expanded Vng & Ilg of a 480V drive @ 30 Hz, 300' of motor cable
[ "0" at offest 4 div; 50 µs/div; 2 A/div; 100V/div]
0-7803-4943-1/98/$10.00 (c) 1998 IEEE
AC Drive
R
+
Iao
XO
PE
450
A+
S
Iao
Motor Frame
U
Iao
V
A-
T
W
CSlot
-
Iao
PE
Logic
Iao Some HF
Potential #1
Motor
Capacitance
350
PE Ground
If Required
By Code
300
PE
Iao
Common Mode
Current
400
Motor
Accidental
Contact Of
Conduit
Iao
Motor PE
Potential #2
Line to Line
250
200
Potential #3
150
Line to Conduit Ground
100
TE Potential 4
50
Fig. 8. System path of lIg zero sequence current
0
100
1000
10000
100000
1000000
10000000
Frequency [ Hz ]
Fig. 12. Measured Zlg & Zll surge impedance's of three #12 awg PVC
wires and a PVC insulated PE wire in a grounded conduit
8 Apk Ilg of Conduit
Fig. 9. Phase current to one hp cable showing Ill & Ilg transient current spikes
100
Normal Mode Resistance
Line to Line
1 pu R =40 mohm
1 pu L = 3.75 uH
1 pu C = 40 pf
10
1 Apk Ilg of PE Wire
Normal Mode Inductance
Line to Line
Vng step
1
Fig. 13. Ilg of Conduit & Ilg of PE ground wire during Vng step change
[ "0" at offset 3 div; 5 µs/div ; 2 A/div; 2 A/div; 100V/div]
Normal Mode Capacitance
Line to Line
0.1
100
1000
10000
100000
1000000
10000000
8
Frequency [ Hz ]
Fig. 10. Measured normal mode R-L-C characteristics of three #12 awg
PVC wires and a PVC insulated PE wire in a 10’grounded conduit
Transformer ( X/R ) Ratio
7
6
Transformer % Z Rating
100
5
1 pu R = 80 mohm
1 pu L = 4.75 uH
1 pu C = 250 pf
4
Common Mode Resistance
Wire to Conduit
3
10
2
1
Common Mode Capacitance
Wire to Conduit
1
0
100
600
1100
1600
Transformer kVA Rating
Common Mode Inductance
Wire to Conduit
Fig. 14. Estimated transformer (% Z) and (X/R) Ratio vs. kVA rating
0.1
100
1000
10000
100000
1000000
10000000
Frequency [Hz]
Fig. 11. Measured common mode R-L-C characteristics of three #12 awg
PVC wires and a PVC insulated PE wire in a 10’ grounded conduit
0-7803-4943-1/98/$10.00 (c) 1998 IEEE
Lcommon
mode
Rext
Rcore
Lcoil
Rcoil
Rext
Rcore
Lcoil
Rcoil
Rext
Fig. 19a. Ilg simulation results for a Vng step input into 163 cables at time
zero with the 2.5 MVA and 250 kVA transformer, input conduits, output
filter and cable skin effect resistors R7–R8 all removed from Fig. 18 circuit
path. [500 A/div ; 1µs/div]
Rcore
Lcoil
Rcoil
Fig. 15. Special PWM output filter with iron core inductor, external damping
resistor and common mode inductor of high frequency ferrite material.
1.2
1 pu phase inductance = 80 uh
1
0.8
0.6
0.4
0.2
0
100
1000
10000
100000
Frequency [ Hz ]
1000000
10000000
Fig. 19b. Ilg simulation results for a Vng step input into 163 cables at time
zero with the 2.5 MVA and 250 kVA transformer, input conduits and output
filter all removed from Fig. 18 circuit path but with incorporating cable skin
effect resistors R7 – R8. [100 A/div; 1 µs/div]
Fig. 16. Inductance vs. frequency of the iron core reactor used in the PWM
filter. The inductance is only 18 % of the 60 Hz value during switching.
1.0E-06
1.0E-07
1.0E-08
1.0E-09
Fig. 19c. Ilg simulation results for a Vng step input into 163 cables at time
zero with only the 250 kVA transformer and output filter removed from Fig.
18 circuit path. [100 A/div; 20 µs/div]
1.0E-10
0.1
1
10
100
1000
AC Motor Horsepower Rating [ hp ]
Fig. 17. Measured stator winding to ground capacitance vs. hp rating
Xo
I2
R1
I1
L1
I4
Xo
R2
L2 C1
I3
L3 R4
R3
OUTPUT
FILTER
INPUT
CONDUIT
250 kVA XFMR
2.5 MVA XFMR
L4
R5 C2
I5
R6
V1
IGBT
I7
OUTPUT ONE
CONDUIT MOTOR
I9
L6
R7 C3
L5 I8
C4
I10
L7
COMBINED MODEL OF QTY 162
OUTPUT CONDUITS & MOTORS
R8 C5
C6
Fig. 18. P-SPICE equivalent circuit model of the zero sequence path with all
components defined at fn.
Fig. 20. Ilg simulation results for a Vng step input into 163 cables at time zero
(middle cursor) with only the 250 kVA transformer removed from Fig. 18
circuit path. Also shown is total current into the combined 162 motor’s line
to ground Cmg capacitance or C6 [50 A/div; 50µs/div]
0-7803-4943-1/98/$10.00 (c) 1998 IEEE
Fig. 21. Final solution Ilg simulation results (top trace) for a Vng step input
into 163 cables at time zero (middle cursor) with all components of Fig. 18
circuit path, including the resistive wye ground isolation transformer &
output filter. Also shown is the 250 kVA X0 resistor current (middle trace)
and 2.5 MVA secondary current (bottom trace) [10 A/div; 20 µs/div]
Fig. 22. Ilg simulation results for a Vng step input into 163 cables at time zero
with all final solution components of Fig. 18 circuit path, except the 250
kVA transformer’s wye X0 neutral is solidly grounded. Also shown is 250
kVA transformer primary current (“o” level trace)
[50 A/div; 20 µs/div]
Fig. 23b. Final solution Ilg on site measured results for a Vng step input (top
trace) into 163 cables at time zero (near middle cursor) with all components
of Fig. 18 circuit path, including the resistive wye ground isolation
transformer & output filter. [10 A/div; 20 µs/div]
Fig. 24. Long term view of Fig. 23a conditions showing Ilg on site measured
results (top trace) for 163 cables with all components of Fig. 18 circuit path,
except the 250 kVA isolation transformer with a 250 ohm resistive wye
ground. [50 A/div; 5 ms/div]. Bottom trace is phase current at drive output
frequency of 30 Hz [200 A/div; 5 ms/div]
Fig. 23a. Final solution Ilg on site measured results for a Vng step input (top
trace) into 163 cables at time zero (near middle cursor) with all components
of Fig. 18 circuit path, except the 250 kVA isolation transformer with a 250
ohm resistive wye ground [50 A/div; 50 µs/div]
Fig. 25. Long term view of final solution Fig. 23b conditions showing Vll
waveform switching into 163 cables at 30 Hz vs. the low Ilg on site measured
current in the PE insulated ground wire (bottom trace) [10 A/div; 500V/div
10 ms/div].
0-7803-4943-1/98/$10.00 (c) 1998 IEEE