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IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 54, NO. 1, JANUARY/FEBRUARY 2018
Modeling, Simulation, and Testing of Switching
Surge Transients in Rapid Transit Vehicles
DC Power Systems
Maxime Berger , Student Member, IEEE, Jean-Pierre Magalhaes Grave , Carl Lavertu, Member, IEEE,
Ilhan Kocar, Senior Member, IEEE, Jean Mahseredjian , Fellow, IEEE, and Daniele Ferrara , Member, IEEE
Abstract—During the operation of dc rapid transit systems, the
rail passenger vehicles are subjected to surge transient events that
can damage the on-board equipment and cause service interruption. During the testing phase, vehicle manufacturers must demonstrate the performance of the vehicles under the specified transients. However, system testing is generally not available during
design. Engineering analysis must be performed to ensure that the
design will meet the transient requirements, and that the overvoltage and overcurrent protective devices are coordinated on a system
level. In response to these design challenges, detailed time-domain
simulation models of a switching surge transient generator and the
vehicle equipment have been developed and validated experimentally. These models are used to evaluate vehicle system parameters
sensitivity, as well as to provide design guidelines for increased
vehicle power system safety, reliability, and availability.
Index Terms—dc power systems, electromagnetic transients program (EMTP), power system protection, railway safety, surge protection.
I. INTRODUCTION
APID transit systems are widely used in North America for high volume passenger transportation. They are
composed of rail passenger vehicles operating as multivehicle
trains on a local dc traction power system [1]. Substations are
equipped with 6- and 12-pulse diode rectifiers. The power from
the traction system substations is delivered to the moving trains
by a third-rail system, or an overhead catenary system at voltage
between 600 and 1500 V [2], [3].
During operation, the rail passenger vehicles are facing surge
transient events that can seriously damage the on-board equip-
R
Manuscript received May 25, 2017; revised September 5, 2017; accepted
October 2, 2017. Date of publication October 10, 2017; date of current version
January 18, 2018. Paper 2017-ESafC-0374.R1, presented at the 2017 Electrical
Safety Workshop, Reno, NV, USA, Jan. 28–Feb. 3, and approved for publication
in the IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS by the Electrical Safety
Committee of the IEEE Industry Applications Society. (Corresponding author:
Maxime Berger.)
M. Berger is with Bombardier Transportation, Polytechnique Montreal, Montreal, QC H3T 1J4, Canada (e-mail: maxime.berger@polymtl.ca).
J.-P. M. Grave, C. Lavertu, and D. Ferrara are with Bombardier Transportation, St-Bruno, QC J3V 6E6, Canada (e-mail: jean-pierre.magalhaes@rail.
bombardier.com; carl.lavertu@rail.bombardier.com; danfer80@outlook.com).
J. Mahseredjian and I. Kocar are with Polytechnique Montreal, Montreal,
QC H3T 1J4, Canada (e-mail: jean.mahseredjian@polymtl.ca; ilhan.kocar@
polymtl.ca).
Color versions of one or more of the figures in this paper are available online
at http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TIA.2017.2761870
Fig. 1.
Available magnetic energy in a typical dc rapid transit system.
ment and cause major service interruption [4]–[7]. These surges
can be predictable or unpredictable [8]. In dc rapid transit systems, surge transients are mainly the consequence of vehicle
regenerative braking action [5], lightning strikes [6], [9]–[11],
rectifier ac-side capacitor bank switching [12], and stored magnetic energy release at fault clearing [13]. The waveform and
the energy level of the surge depends on the nature of the surge
transient.
High-energy switching surge results from the release of magnetic energy stored into the dc traction system track at fault
clearing. Fig. 1 shows an evaluation of the magnetic energy
stored as a function of the fault distance from a typical Montreal
subway system substation (750 V, 2.5 MW, 12-pulse rectifier).
This curve is obtained using the developed substation model
in [1]. At fault clearing, if the dc traction system is not receptive, the energy needs to be dissipated by both the wayside and
the vehicle on-board equipment. Maximum energy condition
typically occurs at a close distance from the substation [2].
Surge transient specifications for rail vehicles are generally
based on standardized waveforms, but transit authorities may
specify their own requirements based on the knowledge of their
local traction system. During the testing phase, vehicle manufacturers must demonstrate the performance of the vehicles under
the specified transients [14]. Experiences have shown that the
investigation of protection solutions during the testing phase
is costly [15]. Complete analysis, considering the behavior of
the on-board equipment, should be performed during the design
0093-9994 © 2017 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission.
See http://www.ieee.org/publications standards/publications/rights/index.html for more information.
BERGER et al.: MODELING, SIMULATION, AND TESTING OF SWITCHING SURGE TRANSIENTS
Fig. 2.
823
Typical dc rapid transit vehicle primary power system.
phase to reduce the risk of protection issues and enhance the
system operational performance, such as safety, reliability, and
availability.
Laboratory testing also showed that the behavior of the vehicle under surge transients is not always intuitive. This is
explained by the complex interaction between the propulsion
inverters and auxiliary power supply (APS) LC-filters, the different overcurrent and overvoltage protective devices mechanisms, and the presence of electronic protection schemes. It is
also worth to mention that there is currently no standardized
approach for surge transient analysis in dc rapid transit systems
[13], [16].
This paper contributes to the knowledge in modeling dc rapid
transit vehicles for simulation, and analysis of high-energy
switching surge transients. In Section II, detailed models of
vehicle’s equipment and protective devices are developed. In
Section III, these models are implemented in the Electromagnetic Transients Program (EMTP) [17], and validated experimentally using data gathered during surge transient testing.
Based on the experimental results, guidelines on the required
level of modeling details are provided for accurate simulations.
In Section IV, the developed simulation model is used to evaluate the influence of the main system parameters on the overall
vehicle operational performance under switching surge transients. Design guidelines are provided based on the simulation
results. Important safety related aspects are also discussed.
II. VEHICLE MODELING
A. Overview of the Vehicle Primary Power System
A typical dc rapid transit vehicle on-board primary power
system is shown in Fig. 2.
The propulsion system of modern vehicles is composed of one
or two traction inverter(s) each connected to one or two threephase traction motor(s). A brake chopper system is also used to
dissipate extra braking energy in case of nonreceptivity of the
dc traction system during dynamic braking operation [18]. The
brake chopper can also act as an overvoltage protective device
upon detection of overvoltage conditions. A high-speed circuit
breaker (HSCB) is also generally used to protect the propulsion
system against internal faults.
Fig. 3.
High-speed circuit breaker (HSCB) timing diagram.
The train auxiliary subsystems are supplied by an APS. The
APS converts the dc traction power system high voltage supply
into lower galvanically isolated dc and ac supplies. The auxiliary
fuse is used to protect the APS against internal faults, while surge
arrester or crowbar circuit [16] may be used for overvoltage
protection inside the APS. Input diodes are also used in many
applications to protect against negative voltage events.
Both the propulsion system and the APS are also equipped
with input LC-filter to meet the railway electromagnetic compatibility requirements. Both air- and iron-core inductors are
used.
B. Equipment Models
1) High-Speed Circuit Breaker (HSCB): The HSCB model
has been previously presented in [1] in the context of transient
overcurrent protection in railway vehicles, but it is summarized
here again for the sake of integrity of the presented material.
HSCB are typically defined using first-order circuit approximation [19]. The HSCB model consists of a controlled voltage
source used to replicate the circuit breaker arcing voltage as
shown in Fig. 3. Laboratory testing confirmed that this model is
accurate enough to approximate the energy dissipated into the
arc during current interruption.
In Fig. 3, when the circuit breaker current if exceeds the
current setting Id , a trip signal is sent to the HSCB trip coil by
its internal detection circuit. As seen in [1], the detection time
td decreases with the prospective current Iss and increases with
the circuit time-constant (L/R). An opening delay tm is present
before the breaker contacts start to open. This delay decreases
with the initial rate-of-rise of the current [20].
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IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 54, NO. 1, JANUARY/FEBRUARY 2018
the selection of the pre-arc I 2 t threshold because it depends on
circuit parameters such as the operating voltage, the frequency,
and the L/R ratio. It is also recommended in [24] to add a 25%
margin when using the pre-arc I 2 t energy
(2)
i2 tm es (t) = (im es (t))2 dt.
Fig. 4.
Equivalent circuit of the current-limiting fuse model.
Once the contacts start to open, an arc voltage ua is introduced into the circuit, and is used to rapidly break the dc current. The initial arc voltage is represented by a step voltage Ui
of duration ti . As the distance between the contacts increases,
the arcing voltage ua starts to increase rapidly. This process is
approximated by a ramp function with slope β. The peak letthrough current Ip occurs at time tp . For standard HSCB, the
difference between tp and td is typically smaller than 10 ms
[2], [21]. The arcing voltage ua then continues to increase up
to a maximum value Ua , which is considered constant following manufacturer’s data for a given arc chute [20]. During the
constant arc voltage period ta, the arcing current ia decreases.
The rate-of-decrease of the arcing current depends on the circuit
inductance and the maximum arc voltage value Ua . Once the
arcing current ia reaches zero, the arc extinguishes.
2) Current-Limiting Fuse: The equivalent circuit of the
current-limiting fuse model is shown in Fig. 4 and is detailed in
Fig. 5. For modeling purpose, the current-limiting fuse model
is divided into two sections: the melting energy calculation, and
the arcing mechanism.
For transient analysis by simulation, the melting energy calculation is performed using both the fuse time–current curve
(TCC) and the pre-arc I 2 t data as shown in Fig. 5(a).
Traditional assessment of fuse performance using ac rms TCC
is a limited practice for transient overcurrent analysis in dc circuits. The equivalent rms value of the instantaneous current
needs to be calculated to use the ac rms TCC. This is because
there is as many equivalent TCC in dc circuit as available fault
current waveforms in the system [22]. An interesting representation of the influence of the time-constant on the fuse TCC is
shown in [23]. In the model of Fig. 5(a), the fuse melting time is
first evaluated by comparing the measured rms current with the
fuse ac rms melting TCC, and then by calculating the equivalent
per-unit melting energy using (1). The total melting time is the
time required for (1) to reach a value of 1 p.u. [15]
1
dt.
(1)
T rm (t) =
tm f (t)
Current-limiting fuse TCC is generally given down to a minimum melting time of 10 ms. Even if it is not clearly represented
on the TCC, current-limiting fuse can break higher fault current in less than 10 ms. In this operating region, the melting
time highly depends on the fault current waveform. I 2 t coordination is recommended in this region according to the IEEE
Std C37.48.1 [24]. In the model of Fig. 5(a), this is done by
calculating the measured pre-arc (or melting) I 2 t energy using
(2) and by comparing the obtained value to the pre-arc I 2 t data
[25]. It is important to carefully analyze manufacturer’s data for
Once it is determined that the fuse is melting from the TCC
or the pre-arc I 2 t, the fuse melting process is modeled as a nonlinear equivalent resistance [see Fig. 5(d)] [26]. The resistance
is adjusted to best fit the peak let-through current data of the
fuse manufacturer for standardized waveforms and testing conditions. There is however a need for a standardized approach
to model current-limiting fuse in time-domain tools based on
available manufacturer’s data.
3) Propulsion System and APS: The equivalent model
schematic for the propulsion system and the APS are given
in Fig. 6.
Both systems require a precharge circuit to reduce the inrush
current at the energization of the converter or after running over
a rail gap. This circuit is modeled by two ideal-switch and a
precharge resistance.
Both the traction inverter and the auxiliary converter are modeled as constant power loads behind an equivalent diode modeling the converters insulated gate bipolar transistor (IGBT)
antiparallel diodes. The antiparallel diodes should be modeled
because they can carry current when the voltage is negative
across the filter capacitor. They need to be modeled by their
equivalent nonlinear VI-characteristic to approximate their energy. Details on the constant power load model are provided in
[27].
As mentioned before, both systems are equipped of input LCfilter. Capacitor and inductor equivalent series resistance should
be considered to increase the model accuracy. If saturable-core
inductor is used, it is important to model its saturation because
once saturated, it is less effective in limiting the surge current.
This has been shown to be one of the most important parameters
to consider in the case of rail vehicles. This is demonstrated in
Section IV.
4) Propulsion Brake Chopper: The propulsion brake chopper is also an efficient overvoltage protective device. The simplest brake chopper control logic is shown in Fig. 7. It is based on
a comparator circuit with a hysteresis band. For surge transient
analysis, the IGBT is modeled by an ideal controlled switch.
Calculation of the energy absorbed by the brake chopper during surge transient is approximated by integrating the power
dissipated by the braking resistor over time as follows:
(3)
Ebr (t) = vbr (t)ibr (t) dt.
5) Metal-Oxide Varistor (MOV): MOV are commonly used
for overvoltage protection in dc applications. The maximum
continuous operating voltage (MCOV) of the MOV is first selected based on the nominal voltage, the available temporary
overvoltage in the system, as well as the permissible leakage
current. Once the MCOV is defined, it is important to evaluate
BERGER et al.: MODELING, SIMULATION, AND TESTING OF SWITCHING SURGE TRANSIENTS
Fig. 5.
Fig. 6.
825
Diagram of the current-limiting fuse model (a) melting energy calculation, (b) breaking process, (c) melting TCC, and (d) nonlinear resistance model.
(a) Propulsion system model. (b) Auxiliary power supply model.
the required clamping voltage by the selection of a MOV with
an appropriate VI-characteristic. It is also mandatory to select a
MOV with the necessary energy-handling capacity for the surge
environment [13].
Modeling the MOV VI-characteristic is necessary for insulation coordination and to evaluate the necessary energy-handling
capacity. Minimum and maximum VI-characteristic tolerances
should be considered. The MOV energy is calculated using (4),
and compared to the energy capability provided by the MOV
manufacturers. Once again, it is important to carefully analyze
manufacturer’s data because the energy rating depends on the
testing conditions, which can differ between different manufacturers [28]
(4)
EM OV (t) = vM OV (t)iM OV (t) dt.
III. EXPERIMENTAL VALIDATION OF SIMULATION MODELS
A. General Procedure
Fig. 7.
Brake chopper equivalent circuit model.
The models presented in Section II are implemented in the
EMTP. The EMTP is a time-domain circuit simulation tool
based on the modified augmented nodal analysis. It uses fixed
time step to solve the circuit system of equations [29]. A time
step of 10 μs is selected to accurately represent all the dynamics
involved in the switching surge transients under study.
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Fig. 8.
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 54, NO. 1, JANUARY/FEBRUARY 2018
Simplified schematic of the surge generator.
Fig. 10.
Validation of the propulsion system and HSCB models (+3 p.u. test).
Fig. 11.
HSCB arcing and brake chopper clamping action (+3 p.u. test).
Fig. 9. Simplified schematic of the surge transient tests. (a) Propulsion system
and HSCB (b) Auxiliary power supply (APS).
A model of an in-house switching surge transient generator
has been developed and implemented in the EMTP (see Fig. 8).
The model of the propulsion system, including the HSCB and
the brake chopper, is first validated with a +3 p.u., 60 kJ surge
injection [see Fig. 9(a)]. The model of the APS including its
current-limiting fuse is then validated with ±3 p.u., 60 kJ surge
injections [see Fig. 9(b)]. Specific conditions related to each of
the three tests are provided in Sections III-C and III-D.
B. Modeling of the Surge Generator
For the validation of the vehicle model, a custom surge transient generator has been used (see Fig. 8). It is composed of
a nonreceptive substation rectifier and a series injection surge
transient source. It allows to perform +/−3 kV, 60 kJ switching
surge transient tests. The surge transient generator power components and the control circuitry are replicated in the simulation
model for increased precision.
C. Validation of the Propulsion System Model
The model of the propulsion system including both the HSCB
and the rheostatic brake chopper is first validated experimentally [see Fig. 9(a)]. A surge of approximately +3 p.u., and
60 kJ is applied to the propulsion system while running at full
power (∼700 kW). The propulsion system has an air-core inductor. The HSCB overcurrent trip setting is set at 2.2 p.u. The test
results are compared with the simulation results in Fig. 10. Both
the surge voltage and the surge current are in good agreements.
In the simulation results of Fig. 11, since the HSCB current
exceeds its trip setting, the HSCB trips and part of the surge
energy is dissipated into the HSCB arc. The brake chopper is
also dissipating part of the surge energy to control the capacitor
voltage between its hysteresis band (1.30–1.35 p.u.). The remaining energy is absorbed by the traction motors and the filter
capacitor.
D. Validation of the APS Model
The model of the APS is also validated experimentally [see
Fig. 9(b)]. The specific APS under test does not have an input
diode but do have a saturable-core inductor. A surge of +3 p.u.,
60 kJ is still injected into the APS while it is lightly loaded. The
test results are also successfully compared with the simulation
results in Fig. 12.
From Fig. 12, it is concluded that the filter inductor saturation
is an important contributing factor because the rate-of-rise of the
surge current increases drastically once the inductor saturates.
It can also be seen that the MOV is clamping the surge voltage
BERGER et al.: MODELING, SIMULATION, AND TESTING OF SWITCHING SURGE TRANSIENTS
Fig. 12.
Fig. 13.
Fig. 15.
Impact of the HSCB on the propulsion units surge current.
Fig. 16.
Impact of the HSCB on the capacitor voltage.
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Validation of the APS model (+3 p.u. test).
Calculation of the energy absorbed by the MOV (+3 p.u. test).
these diodes and evaluation of their energy is mandatory to
ensure that they will not be damaged.
It must also be mentioned that, in both cases, the melting
energy of the fuse, which is calculated using (1) and (2), is not
high enough to break the fuse.
By looking at Figs. 10, 12, and 14, it is also obvious that
the simulation models are accurate enough to model both the
propulsion system and the APS. Details such as inductor saturation, MOV VI-characteristic, fuse melting energy, HSCB detection and arcing mechanisms, brake chopper control scheme,
and IGBT antiparallel diodes must be considered for accurate
simulation and analysis of the entire system behavior.
IV. ANALYSIS OF MAIN PARAMETERS INFLUENCE
Fig. 14.
Validation of the APS model (–3 p.u. test).
to 1.9 p.u. at the input terminal of the APS unit. In Fig. 13, the
energy absorbed by the MOV is calculated by simulation to be
49 kJ using (4).
A negative test is also applied to the APS to demonstrate the
importance of modeling the IGBT antiparallel diodes.
The results from Fig. 14 show that the filter capacitor voltage
is at some point clamped to zero because the IGBT antiparallel
diodes are conducting once there is a polarity reversal at the
capacitor terminals. A large amount of energy may pass through
A generic vehicle surge transient test bench is implemented
using the validated models. It is used to provide a better understanding of the system behavior under surge transients. Parameters sensitivity will be discussed throughout different case study.
The main simulation parameters are provided in the Appendix.
The vehicle is composed of two propulsion system units, and
one APS. The loads are neglected to observe capacitor trapped
charge.
A. HSCB Detection and Arcing
A surge transient of +2.25 kV, 50 kJ is applied into the
propulsion system. Under surge transients, the propulsion system can benefit from the use of the HSCB if the surge current
exceeds its trip setting (see Fig. 15). The HSCB can in fact reduce the energy transferred into the propulsion system, which
can further reduce the level of trapped charge at the filter capacitor. HSCB tripping under high surge transient is acceptable from
an operational standpoint because it can be reclosed remotely.
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Fig. 17.
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 54, NO. 1, JANUARY/FEBRUARY 2018
Fig. 18.
Impact of the inductor saturation on the APS unit.
Fig. 19.
Impact of the input diode on the APS antiparallel diode current.
Impact of the MOV on fuse melting energy and capacitor voltage.
As shown in Fig. 16, the HSCB can reduce the level of trapped
charge at the filter capacitor. These trapped charges are due to
limitations with simple hysteresis band control of the brake
chopper. Under running condition, these trapped charges are
dissipated by the traction motors. Under the no-load condition,
these trapped charges may take time to discharge. This is important to be considered when selecting the brake chopper turn-on
and turn-off thresholds.
B. Current-Limiting Fuse and MOV Coordination
Coordinating current-limiting fuse with MOV is important.
Historically, surge arresters have been placed on the load side
of the protection fuse for maintainability reasons associated
with the localization of the equipment boxes. The disadvantage
of this configuration is that the MOV acts as a low resistance
during the surge transient and may lead to nuisance operation
of the fuse [24]. Fuse should be avoided to break during surge
transient. The surge energy should be absorbed by the MOV
without breaking the fuse to allow the system to quickly recover
after the event.
As an example, a surge transient of +3 kV, 60 kJ is applied
into the APS system. Fig. 17 shows the melting energy (in p.u.)
calculated with the fuse TCC using (1). It also shows the melting
energy in kA2 s calculated using (2) which is compared with the
I 2 t melting energy provided by the fuse manufacturer. During
surge transient, Fig. 17 shows that a MOV located on the load
side of the fuse increases the melting energy seen by the fuse,
but it also reduces drastically the level of trapped charge at the
filter capacitor.
C. Inductor Saturation
In railway systems, second-order input LC-filters are typically
used. Typically, at frequencies higher than 50/60 Hz, the input
filter must be inductive. Cut-off frequencies in the range of 15–
35 Hz are generally selected, and the values of the capacitor and
inductor are selected accordingly.
It is important to design the input filter to meet the electromagnetic compatibility requirements including all the surge
transients. In addition, the converter design must consider the
influence of the input filter on the converter transfer functions
[30].
The impact of saturable-core inductor on the ability of the
filter to protect the APS unit against surge transients is shown in
Fig. 18. For the same surge applied to the APS (+3 kV, 60 kJ),
the air-core inductor is more effective than the saturable-core to
limit the surge current rate-of-rise, which leads to a reduction
of the maximum capacitor voltage. Air-core inductor should
be preferred when high level of switching surge transients are
expected.
D. APS Input Diode
Input diode cannot be used in the propulsion system because
regenerative braking requires the power to flow in both directions at the converter terminals. However, for the APS, input
diode can enhance the converter ability to deal with negative
voltage surges.
As an example, a –3 kV, 60 kJ surge is applied into the complete system (two propulsion units and one APS). The results are
shown in Fig. 19. The input diode adequately protects the IGBT
antiparallel diodes during the surge transient. However, when
the input diode is not present, the negative surge propagates at
the capacitor terminals, which leads to a polarity reversal. Since
the IGBT antiparallel diodes are also in parallel with the filter
capacitor, they carry part of the surge current. When there is
no input diode, it is recommended to make sure that the IGBT
antiparallel diodes I 2 t capability is not exceeded.
BERGER et al.: MODELING, SIMULATION, AND TESTING OF SWITCHING SURGE TRANSIENTS
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V. CONCLUSION
Fig. 20.
Current-limiting fuse TCC implemented in the EMTP.
Overall, this paper contributes to the knowledge in modeling
dc rapid transit vehicles for simulation, and analysis of highenergy switching surge transients. Detailed models of a surge
transient laboratory setup including the vehicle equipment under test and the protective devices have been developed and
validated experimentally.
Using the developed and validated vehicle models, this paper identifies important modeling considerations for accurate
analysis of surge transients in dc rapid transit vehicles. Through
different case study, it also presents important design guidelines which can also be extended for any converter designed for
operating in a similar surge environment.
One important conclusion of this work is that vehicles should
not be excluded of surge transient analysis in dc rapid transit
systems. They are both part of the problems and the solutions,
i.e., fault inside a vehicle can generate surge conditions, but the
on-board protective devices are also a mean of absorbing surge
energy circulating into the dc traction system.
Experiences have also shown that the development of accurate models is a crucial step to reduce design iterations. Once
the models are validated, the analysis of the system behavior
is easier using simulations than by high-power laboratory testing. A good knowledge of the system behavior is mandatory to
improve system safety, reliability, and availability.
APPENDIX
MAIN SIMULATION PARAMETERS
Fig. 21.
MOV VI-characteristic implemented in the EMTP.
Model
E. Summary of the Lessons Learned on Safety Aspects
In rapid transit application, passengers’ safety and vehicle’s
essential systems availability are closely related. It is important
to keep supplying power to the propulsion system and the auxiliary power system to reduce the occurrence of train failure
during operation. On one side, during emergency conditions,
availability of the propulsion system allows to move the train
to the next station for safe evacuation of the passengers. On the
other side, equipment such as lighting, communication devices,
ventilation system, and door control mechanisms are critical to
avoid a general panic of the passengers, and safely evacuate the
train during emergency conditions [27].
In this context, an appropriate coordination of the overvoltage and overcurrent protective devices is mandatory to increase
the availability of the important systems. The utilization of the
air-core filter inductor, and input diode for the auxiliary power
system also increase system resiliency to surge transient events.
These are important conclusions for increasing system availability, and, therefore, passengers’ safety.
Furthermore, appropriate selection of protective devices ratings is necessary to avoid failures leading to hazardous conditions. A good knowledge of the surge environment and the
limits of the protective devices (e.g., energy-handling capacity
of MOV) is also critical to reduce the risk of protective device
failure.
Description
Symbol
Value
–
–
750 V
900 V
Traction System
Nominal Voltage
Maximum Voltage
Propulsion
System
(2 units)
Power Set-Point
Input Filter Inductor
Input Filter Capacitor
–
Lf
Cf
No-Load
2.5 mH
8.2 mF
Brake
Chopper
(2 units)
Turn-on Voltage
Turn-off Voltage
Brake Resistor
–
–
Rb r
1010 V
975 V
0.3 Ω
Detection
Id
2 kA
Arc Voltage
Ui
Ua
tm
ti
β
50 V
1500 V
4 ms
2 ms
480 V/ms
HSCB
Auxiliary Fuse
Time–Current Curve
Melting Energy
–
–
See Fig. 20
140 kA2 s
MOV
VI-Characteristic
–
See Fig. 21
APS
Power Set-Point
Input Filter Inductor
–
Lf
Input Filter Capacitor
Cf
No-Load
6.4 mH (@ 0 A)
1.7 mH (@ 275 A)
5.6 mF
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Maxime Berger (S’14) received the B.Eng. degree
from the Université du Québec à Rimouski, Rimouski, QC, Canada, in 2014, and the M.A.Sc. degree in electrical engineering from Polytechnique
Montréal, Montréal, QC, in 2016, where he is currently working toward the Ph.D. degree in electrical
engineering.
Since 2012, he has been with Bombardier Transportation, St.-Bruno, QC, where he has been working
in the Systems Integration Department. His research
interests include power electronics, protection, and
the simulation and analysis of power systems transients. His research is supported by Bombardier Transportation, the Fonds de Recherche du QuébecNature et Technologies, and the Natural Sciences and Engineering Research
Council of Canada through the Industrial Innovation Scholarship Program.
Jean-Pierre Magalhaes Grave received the B.Eng.
degree in electrical engineering from École Polytechnique de Montréal, Montréal, QC, Canada, in 2004.
From 2004 to 2011, he was at Oerlikon Transtec,
Inc., St.-Jean-sur-Richelieu, QC. Since 2011, he has
been with Bombardier Transportation, St.-Bruno,
QC, where he is currently an Expert in auxiliary systems. He is involved in the design, integration, and
testing of auxiliary power converters, battery systems,
and lighting systems for rail passenger applications.
He has two U.S. Patents.
Mr. Magalhaes is a registered Professional Engineer in the province of QC.
Carl Lavertu (M’17) received the B.Eng. degree in
electrical engineering from the Université du Québec
à Trois-Rivières, Trois-Rivières, QC, Canada, in
1985.
From 1985 to 1991, he was at Lauzon Lavertu
Consultants, Inc., Drummondville, QC. Since 1991,
he has been with Bombardier Transportation, St.Bruno, QC, where he is currently a Senior Expert in
power electronics. He has been working on various
rail projects, being mainly involved in the design, fabrication, and testing of traction and auxiliary power
converters, battery systems, door controllers, and lighting systems. He has two
U.S. Patents.
Mr. Lavertu is a registered Professional Engineer in the province of QC.
Ilhan Kocar (M’10–SM’13) received the Ph.D. degree in electrical engineering from the École Polytechnique de Montréal, affiliated with Université de
Montréal, Montréal, QC, Canada, in 2009.
He was a Project Engineer at Aselsan Electronics, Inc., St.-Bruno, QC, Canada from 1998 to 2004.
He was a Research and Development Engineer at
CYME International T&D, St.-Bruno, QC, Canada
from 2009 to 2011. In 2011, he joined the Faculty
of Electrical Engineering with Polytechnique Montreal, Montréal. He works in close collaboration with
electric utilities, research organizations, and power system tool developers. His
research interests include power system analysis, modeling, and simulation.
Dr. Kocar is a registered Engineer in the province of QC.
BERGER et al.: MODELING, SIMULATION, AND TESTING OF SWITCHING SURGE TRANSIENTS
Jean Mahseredjian (F’13) received the Ph.D. degree
in electrical engineering from École Polytechnique de
Montréal, Montréal, QC, Canada, in 1991.
From 1987 to 2004, he was at IREQ (HydroQuébec), QC, working on research and development
activities related to the simulation and analysis of
electromagnetic transients. In December 2004, he
joined the Faculty of Electrical Engineering with
École Polytechnique de Montréal.
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Daniele Ferrara (M’16) received the B.S. degree
in mathematics for engineering from Politecnico di
Torino, Turin, Italy, in 2002.
From 2006 to 2011, he was at various engineering firms delivering consultancy services for the
aerospace industry in the Turin area. From 2011 to
2017, he held various engineering management roles
at Bombardier Transportation, Vado Ligure, Italy, and
St.-Bruno, QC, Canada. Since 2017, he has been
with SNC-Lavalin, Montréal, QC, where he is currently the Director, Rail Control Systems. He has
been working on various aerospace and rail projects, being mainly involved in
the design, functional integration, and testing of different on-board and wayside
systems.