![Characteristic Impedance equivalent model of considered MVDC link. placed in series with voltage sources to avoid unnecessary reflection from DC source, as depicted in Fig. 9. Thus, impact of the emulated DC voltage source is negligible, and the cumulative characteristic impedance of the converter station is not impacted significantly. In addition, the reflection coefficient of the substation busbar & transmission line tends to reflect 26.06% of substation intruding surge. While the remainder of it would be refracted into the MMC substation. Reflection & transmission coefficient between substation busbar and OHPC are shown in (12) & (13) respectively [37].](https://www.researchgate.net/publication/372264384/figure/fig2/AS:11431281173956346@1689083331362/Characteristic-Impedance-equivalent-model-of-considered-MVDC-link-placed-in-series-with_Q320.jpg)
Available via license: CC BY 4.0
Content may be subject to copyright.
VOLUME XX, 2017 1
Date of publication xxxx 00, 0000, date of current version xxxx 00, 0000.
Digital Object Identifier 10.1109/ACCESS.2022.Doi Number
Analysis of Lightning Transient characteristics
of short-length mixed MMC-MVDC transmission
system
MUHAMMAD USMAN 1, KYU-HOON PARK 1, (Student Member, IEEE), and BANG-WOOK
LEE 2, (Senior Member, IEEE)
1HVDC Electric Power Lab, Department of Electrical and Electronic Engineering, Hanyang University, Ansan 15588, South Korea
2Division of Electrical Engineering, Hanyang University, Ansan 15588, South Korea
Corresponding author: Bang-Wook Lee (bangwook@hanyang.ac.kr).
This work was supported in part by the National Research Foundation of Korea (NRF) and in part by the Ministry of Education (MOE, Korea), South
Korea, under BK21-4 project.
ABSTRACT Medium voltage direct current (MVDC) transmission system are growing due to their
assistive quality in conventional grid and compatibility with renewable power network. MVDC distribution
links with “Mixed” overhead (OH) & underground (UG) sections could be devised based on urban
planning. UG Cables or substations are indirectly exposed to lightning strikes due to adjacent tower
sections. In case of MVDC converter or cable, present researchers do not specify lightning voltage impulse
level for related system voltage. Therefore, preluding electromagnetic (EM) transient investigation are
required to determine the maximum lightning overvoltages for system peripherals i.e. cable & Modular
Multilevel Converter (MMC) substation. This research focuses on analyzing lightning performance of OH
transmission towers-cable junction & tower-substation link in case of a shielding failure (SF) and back
flashover (BF) for a ±35kV short-length mixed MMC-MVDC transmission scheme. This article provides
broad-band modeling method for MMC substation for lightning investigation. In addition, based on a
detailed time-domain parametric evaluation in PSCAD/EMTDC program, lightning impulse voltage across
the transmission line’s pole insulator and embedded cable section are estimated along with numerical
validation relying on travelling wave theory. Effect of project parameters such as tower grounding
resistance, riser section surge impedance (which connects cable & OH line) and cable length on lightning
overvoltage impacting the cable and connected tower section has been demonstrated.
INDEX TERMS MVDC transmission, Modular Multilevel Converter, maximum shielding failure,
insulation flashover, PSCAD/EMTDC, sheath grounding, riser section, DC cable.
I. INTRODUCTION
Commercial application of DC transmission systems has
increased drastically over the past few decades. As
compared to AC power systems, electronic power
converters offer better integration with unconventional/low
footprint power resources and more efficient power
management. This has led to the construction of medium
voltage DC links by energy subsidies and research
consortiums across the globe [1,2] i.e., ANGLE DC project
in Europe, ±10 kV MVDC distribution project in Zhangbei,
Zhuhai and Guizhou and HVDC Light (Denmark-Sweden).
As reported, most MVDC projects are point-to-point links
or based on underground (UG) power cables [2,3]. However,
recent proposals for the conversion of MVAC transmission
lines to DC have paved the development of mixed MVDC
lines [3,4]. These transmission corridors could pass through
urban or suburban regions. Where availability of space,
visual impact, and city planning could influence transmission
infrastructure. Thus, MVDC projects with partial overhead
(OH) and UG transmission segments would be constructed
worldwide like HVDC projects. In such a scenario UG cable
and MVDC substation are exposed to indirect lightning
strike. Overvoltages caused by such events are important to
be investigated for insulation coordination of cable system &
substation. Consequently, lightning overvoltage on cable-
overhead line (OHL) junction has been studied for MVAC
system [5]. Methods to calculate lightning overvoltage surge
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
2 VOLUME XX, 2017
along a DC cable in a mixed transmission system have been
demonstrated in few researches [6,7].
Currently, significant lightning research has been
conducted on “Line Commutated Converter” (LCC) DC
transmission lines at HVDC level including “Mixed”
transmission Links [8]. Contrary to MVDC projects which
are 10-100km in length, HVDC systems have long-length
transmission lines thus reflection and refractions from the
converter substation are not taken into consideration for
lightning studies. Usually, insulation coordination of few
towers and cable sections with varying length are studied [9].
Latest MVDC corridors are composed of modular multilevel
converter (MMC) along with DC specialized switchgear and
DC circuit breakers (CB); standards for such power projects
are being evolved i.e., PROMOTioN Project in Europe.
Some research articles had shed light on high-frequency
MMC lightning study. Gao et al. [10] depicts MMC
converter station by representing the electronic switches as
open/closed switches depending on their state. A detailed
model of submodule (which is building block of MMC
converter) has been defined in reference [11]. This is done by
representing stray capacitance and inductances in an IGBT of
submodule. Zhu et al. has done a thorough electromagnetic
investigation (EMI) of MMC converter station by
representing the structural capacitances of submodules
housed in a vertical configuration [12].
The scope of this paper is to derive lightning impulse
voltage across the UG Cable, adjacent overhead transmission
line (OHTL) tower’s as well as DC substation in a mixed
MMC-MVDC grid. In order to obtain a general statement
about the occurring voltage stresses, parametric evaluation
and time-domain analysis has been performed. A variety of
system parameters like transmission tower grounding
conditions, cable length & riser surge impedance is evaluated
under shielding failure and back flashover lightning current.
It is not intended to conduct entire insulation coordination
studies within this research. However, compared to [5]
understanding and concept about the lightning surge
waveform across cable and transitioning tower insulators
have been broadened for MVDC system. Additionally,
modelling depth and scope regarding wide-band MMC
converter & substation for lightning study has been
fundamentally extended from previous research [10].
This article is further structured as follows. Section Ⅱ
provides a description of ±35kV Mixed symmetric
monopolar MMC-VSC transmission with a brief explanation
of wide-band converter station & transmission line model for
lightning investigation in PSCAD/EMTDC program. Project
specifications for transmission infrastructure from articles
examining MVAC to MVDC distribution system are utilized
and interpolated [7,14]. Section Ⅲ provides an overview of
shielding failure (SF) and backflash-over (BF) lightning
intrusion waveform for system under considered.
Subsequently, a method to include steady state DC system
voltage at the converter station for accurate lightning analysis
is also reported. Section Ⅳ portrays a scenario in which
lightning strike at a tower adjacent to cable connected
transmission tower. Peak lightning transient overvoltage at
various location of OHTL, riser and cable section are
estimated in PSCAD. Furthermore, equations are formulated
to validate the following contribution:
Recognized and explained the superior flashover
performance of transitioning tower under the influence
of SF.
Identified the role of tower footing resistance on farthest
tower insulators’ flashover.
Investigated the variance in BF overvoltage waveform at
tower insulator w.r.t pole’s polarity and explained higher
flashover probability of -ve/+ve pole insulator at
transitioning and adjacent towers as compared to SF.
The highest lightning peak intrusion voltage across the
cable length has been estimated with regards to location
& variation in length.
Role of riser section’s (which connects OH line & cable)
surge impedance on lightning surge on transitioning
tower have been analysed.
Analysed the retarded surge waveform across the cable
when surge arresters are installed.
Section Ⅴ present lightning transient voltages across MMC
converter & DC substation with following contributions:
Analyzed the impact of length & mutual surge
impedance of substation busbar on overvoltage’s at
converter station peripherals in case of SF & BF.
Proved that hypothetical DC voltage source introduced
in section Ⅲ appropriately adds the steady-state system
voltage in PSCAD/EMTDC for lightning studies.
Finally, section Ⅵ & Ⅶ provide a general evaluation and
conclusion respectively.
II. SYSTEM MODELLING AND DESCRIPTION
A. TRANSMISSION LINK PARAMETERS
A half-bridge MMC converter Monopolar ±35kV point to
point link is considered for the manuscript which is
connected to a 22.9kV AC system on both sides. Two 10km
FIGURE 1. A general model of mixed MVDC point-to-point link with VSC
converter station.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
VOLUME XX, 2017 3
TABLE Ⅰ
PARAMETERS OF ±35KV MVDC TRANSMISSION LINK
Equipment
Converter station
1st (Master control) & 2nd (PQ control)
converter topology
Half-bridge sub-module based symmetrical
monopolar configuration
power rating
34.2 MW
DC rated voltage
±35 kV
submodules per arm
28
voltage of submodule
2.5kV (max)
submodule capacitor
3.26 mF
DC nominal current
434 A
bridge arm reactor
17 mH
pole reactor
10 mH
OHL section are connected with underground cable section
in PSCAD model. Where sheath of the cable is solidly
bonded to OHGW. Cable length is taken as variable in this
study to investigate the worst-case scenario. The MMC
converter station is composed of 6 arms, each comprising of
28 half-bridges. The detail of fundamental components of
converter stations is given in Table. Ⅰ along with the basic
structure of the transmission structure, shown in Fig. 1.
Based on the relatively small size of the MVDC converter
station as compared to higher voltage levels, the substation is
considered to be confined indoors. As switchgear for DC
systems are being developed isolated till now i.e., DC circuit
breaker topologies devised are independently built or other
disconnectors/earthing equipment are incorporated according
to converter topology [15,16]. For a short-length
transmission line, one DC interrupter on each pole could be
enough for current fault protection. Thus, a DCCB is
installed on positive pole of one substation & negative pole
of the other substation.
B. CONVERTER STATION MODELLING
To evaluate the response of MMC substation under high-
frequency surge, each part must be modelled appropriately.
For steady state study, converter station sub-modules are
represented as a voltage source with equivalent resistance
[18,19]. For any instance, during normal operation, the
capacitor voltage in submodules face insignificant change as
shown in Fig. 2.
Additionally, lightning surges have a very high frequency
FIGURE 2. Prospect of voltage across the submodule capacitance in an
arm of MMC converter.
FIGURE 3. (a) High-frequency converter station infrastructure layout. (b) Basic electrical configuration of half-bridge submodule. (c) compactly
enclosed MMC submodule. (d) high-frequency equivalent of internal Submodule circuit.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
4 VOLUME XX, 2017
TABLE Ⅱ
DCCB ESTIMATED VALUE OF DIFFERENT ELEMENTS. [25]
R
C
L
RVI
CVI
LVI
Terminal to earth
capacitance (Cg)
Surge arrester
Clamping voltage (SA)
Surge arrester stray
capacitance (CSA)
50mΩ
2.5µF
19µH
50Ω
0.2nF
50nH
15pF
52kV
5.2nF
(more than 10MHz), much large than the switching
frequency of converter station. Computation for such a small
time period doesn’t require conventional MMC converter
models in PSCAD. Thus, voltage source in PSCAD with
converter surge impedance is used to represent MMC for this
investigation.
The MMC converter station arms are taken to be a vertical
pile of connected submodules. Each submodule (SM) is in a
half-bridge configuration with 2 IGBT (1200G450350) [20].
The stray capacitance & inductance of them are used to
model the SM [11]. Although the submodule tower structure
has no electrical connection with the MMC power-electronic
equipment, for high surge overvoltage evaluation, the stray
capacitance of the metallic MMC tower must be accounted.
Zhu et al. [12] discuss a set of stray capacitances subjected to
converter station. The housing capacitance between the
submodules could be convolved into just two types i.e.,
terminal to ground and SM parallel stray capacitance (Fig. 3).
For instance, C(n-27) and C(n-28) are the terminal to ground
stray capacitance of sub-module 28, while C28 represent stray
capacitor across the SM. For such structures, stray elements
must be calculated exclusively. Here, 10pF is utilized for
both types of SM’s stray capacitances. Thus, accounting for
internal & external stray elements of SM’s.
C. SWITCHGEAR
The use of special type of DC interrupter in an MVDC grid is
eminent for grid protection. A lot of DC CB have been
proposed and evaluated but only few of them have been
physically developed and tested i.e. offshore HVDC system
on Nan’ao island in China [24]. A ±27kV Forced Oscillating
interrupter has been developed in the PROMOTioN project
[25]. The stray components of Vacuum interrupter (VI) and
other electronic parts of that DC CB has been considered for
the supposed substation. Detailed information about
additional elements of DC breaker is presented in Table. Ⅱ
and Fig. 4. VI is in closed condition with a resistance of
80µΩ. Wideband modelling of DC CB’s surge arrester is
done as explained in section Ⅵ-F.
Disconnector or breakers are required for energizing and
protection of any general power network. The number and
type of switchgear in a practical MVDC link could vary.
Thus, derived configuration of substation is utilized [15].
Disconnecting switches and measurement transformers are
represented as their parasitic ground capacitances [21-24] (as
depicted in Table. Ⅲ). Necessary elements like Converter
Disconnector switch (CD), Pole Line Earthing switch (PLES)
or Electrode Line Disconnector switch (ELD), etc. are
incorporated (Fig. 5(a)). Typical Air Insulated Substation
(AIS) busbar has a self-surge impedance of 350Ω [29]. To
consider the effect of mutual surge impedance between
busbars, they are represented as frequency-dependent line
model in PSCAD/EMTDC [21]. Dimensions of busbar are
presented in Fig. 5(b). The total length of AIS busbar is 25m
per pole. Additionally, Surge retardation of 20% has been
added to account the presence of bushings, supporting
insulators & measurement equipment in substation. Finally,
the interconnection between the transmission line and
substation are modelled as lumped inductance of 4m
(1µH/m) [9].
D. TOWER AND LINE STRUCTURE
For this study, experimented AC transmission equipment for
DC compatibility is chosen [4]. Compact tower structure
reduces carbon footprint and electrical interference in line
[13]. However, audible corona and insulator flashover limits
the tower compactness. For ±35kV line, the Electric field
strength model of Austrian 30kV AC tower has been
interpolated [4] (Fig. 6(b)), similar tower model has been
utilized in other studies [26]. The considered tower is a
conical ‘T’ shaped galvanized steel pole. The tower is
grounded using a 2m lead wire with a 0.05m radius. Also, a
ground wire is placed 1.545m above the pole conductors.
In PSCAD/EMTDC Frequency-Dependent phase model
has been incorporated that can represent transmission lines
over a wide range of frequencies with a DC Correction
factor. The line is divided into multiple sections in PSCAD to
emulate the overvoltages midway between it. Transmission
tower is simulated as Bergeron line model to represent tower
surge impedance & cross arm/braces surge retardation.
Tower structure’s surge impedance is evaluated
based on its geometry. (1) & (2) are incorporated in PSCAD
FIGURE 4. DC circuit breaker wideband model
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
VOLUME XX, 2017 5
FIGURE 5. (a) MVDC substation depicting the electronic converter and switchgear. (b) Configuration and dimensions of air insulated busbar at
substation.
as transmission line model with 15% surge retardation [27].
Where, Ravg (2) is a ratio composed of height from base to
midspan of tower h1 (m) and midspan to top h2 (m), 4.5m
each, while r1, r2 & r3 are the top, midsection and bottom
radius of the tower structure which are 0.1m, 0.25m and
0.5m repeatedly for 9m conical tower structure [27].
calculated by this scheme is equal to 209 Ω. Short length
grounding rod and tower portion between the OH power
cable (PC) & ground wire (GW) are presented as RLC
lumped model (for earth resistivity ρ 100 Ω.m) and
inductance respectively as shown in Fig. 6(a). [22,28].
E. INSULATOR MODELLING
Pin type ceramic insulator VHD 35-G based on Austrian
standards have been utilized for this study [30], DC
evaluation of similar insulators has been done before [4]. For
the insulator under consideration, the overall capacitance is
100pF which is installed in parallel with the ideal switch
[31]. The inductance of the insulator path can be modelled as
a lumped inductance (1µH/m) in series with the switch as
depicted in Fig. 6(a). Although there are multiple methods to
evaluate insulator flashover under the influence of fast front
transients. However, for insulator of length shorter than 1.2m
‘Disruptive effect method’ could be utilized [32]. This
technique evaluates breakdown process as a function of
voltage applied across the insulator and time duration of the
applied voltage.
TABLE Ⅲ
PSCAD REPRESENTATION OF MMC SWITCHGEAR [10,23]
Equipment
Abbreviation
Capacitance
Converter Disconnector Switch
CD
50 pF (pole
to ground)
Substation Disconnector Switch
SD
Line Disconnector
LD
DC Circuit Breaker
DCCB
Pole Line Disconnector
PLD
Neutral Bus Switch
NBS
Neutral Bus Disconnector
NBD
Electrode Line Disconnector
ELD
Metallic Return Transfer Switch
MRTS
Filter Disconnector
FD
Converter Earthing Switch
CES
Filter Earthing Switch
FES
Substation Earthing Switch
SES
Pole Line Earthing Switch
PLES
Potential transformer
PT
100 pF
Current transformer
CT
50 pF
Smoothing/Pole reactance
--
50 pF
In case, the insulator voltage exceeds a certain value X (kV),
the breakdown of air gap can be evaluated, mathematically
formulated as:
where “D” is Disruption effect/Area criteria specific for a
certain length of insulator. Evaluated by the integral of
difference between instantaneous voltage across the insulator
V(t) and the triggering voltage X, starting from the lightning
triggering instance to. Once the integral value increase above
D the above-mentioned ideal switch is closed in the
simulation to emulate insulator flashover. The constants (X,
K, D) for 0.29m long insulator are given in Table. Ⅳ [17].
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
6 VOLUME XX, 2017
FIGURE 6. (a) High-frequency tower model. (b) Overhead ±35kV
galvanized steel pole with ground wire.
TABLE Ⅳ
DISRUPTION EFFECT MODEL CONSTRAINTS FOR ±35KV INSULATOR [17]
VHD 35-G height (m)
K
X (kV)
D
0.29
0.92
170
0.037
F. UNDERGROUND CABLE
Like PSCAD transmission line representation, the
underground cable could be modelled as frequency-
dependent (phase) model. The cable sections for both the
poles are designed as 1-core XLPE Cable [26], 0.5m apart
from each other and 1.5m deep underground. The core
diameter, insulation, and sheath thickness along with the
single cable depth are depicted in Fig. 7(a). The cable poles
are divided into several sections to evaluate the voltage in
separated sections of the cables. The lightning-withstand-
level (LIWL) of cable core insulator is 170kV [33] (IEC
60071-1) that means if the cable overvoltage, due to
lightning, surpasses this value the inner insulation of cable
may rapture. Similarly, the cable sheath LIWL is 60kV. The
sheath of cable sections can be grounded in multiple ways to
reduce any stress on outer jacket of insulator under transient
conditions [34]. Here, analysis is done by considering multi-
terminal grounding as shown in Fig. 7(b). The sheath is
grounded with 10 Ω (Rg) resistance at each subsection of the
UG transmission section. However, Rg is also varied to study
its impact on cable & sheath overvoltages (section Ⅳ-D).
G. OTHER TRANSMISSION EQUIPMENT
“Mixed” transmission line requires riser section (which is
tower with underground cable connection) between overhead
power cable (OHPC) and underground cable (Fig. 7(c)).
Since there are no proper guidelines for riser section
modelling. Asif et al. [8] has modelled it like an OHL (i.e.,
similar speed of wave propagation, geometry, and surge
impedance). However, practically riser section should have a
geometry transitioning from tower section to cable. For
example, the separation between the poles and ground
conductors reduces gradually and riser sections conductors
must have an appropriate insulation jacket. This suggests that
surge impedance of riser section might be approximately
average of OHPC & underground cables. Thus, the impact of
riser surge variation has been studied in this manuscript.
III. LIGHTNING SURGE INSTIGATION
The lightning surge waveshape impact transients on
transmission equipment. For this study, CIGRE single
lightning stroke is utilized with varying magnitude [22].
Lightning channel impedance of 1000 Ω and 400 Ω (Zc) is
added in PSCAD for SF and back flashover (BF) respectively
as suggested in previous literature [35].
A. LIGHTNING SIGNAL WAVEFORM
In the presence of a ground wire, the lightning strokes
reaching directly to the phase wire are relatively of lower
magnitude. The first lightning stoke can be mathematically
depicted as:
(4)
Where, & form the front portion while &
compose the tail portion of the CIGRE lightning model.
(5)
(6)
While represents the instantaneous time after the initiation
of lightning current. Similarly, & are the front time and
maximum steepness of CIGRE lightning stroke. (see Fig. 8)
(7)
(8)
The maximum shielding failure lightning stroke current
IMSF is calculated by conducting Electro geometric modelling
of the overhead transmission structure (proposed by IEEE
Std. 1243) [36]. The geometric constraint of tower structure,
in this paper, suggests an IMSF of 10kA. Elements of IMSF are
tabulated in Table. Ⅴ and depicted in Fig. 8.
Lightning strokes of higher values could strike at overhead
ground wire (OHGW) causing BF. However, cumulative
probability of very large magnitude is low i.e., chances of
worst-case lightning amplitude of 200kA is less than 1% [36]
which in case of an urban/sub-urban area might not directly
strike a short length MVDC transmission system. In addition,
occurrence of certain lightning impulse along an OHTL is
estimated using project specific ground stroke density. Gao et
al. [10] utilized 118kA BF lightning magnitude estimated for
a period of 20 years based on the length & average ground
lightning density (Ng). For a 35kV mixed AC transmission
line a 150kA BF lightning current is chosen in [5] which is
supposed to be rare. Here, 110kA OHGW lightning have
been considered for BF analysis.
Although, for lightning insulation coordination maximum BF
lightning current should be chosen based on the service life of
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
VOLUME XX, 2017 7
TABLE Ⅴ
PARAMETERS OF CONSIDERED IMSF LIGHTNING CURRENT WAVEFORM.
Maximum steepness
(kA/µs)
Half time
()
()
()
(kA)
(kA)
n
a
(kA/µs)
b
(kA/µs)
17.79
50
65.3369
0.0562
4.7119
10.0086
1.0086
18.0065
0.9763
3.323
FIGURE 7. (a) Detailed cross-sectional data of a XLPE Cable. (b) Multi-terminal sheath grounding of underground transmission segment (c) riser tower
& cable.
the MVDC system peripherals i.e. cable section. However,
the aim of this research is to assess the general BF lightning
transients of the transmission link. Thus, a more frequent and
rounded-off magnitude of 110kA has been utilized for it,
which accounts for the maximum BF occurrence in a 12-year
period for the OHTL tower as shown below.
Average ground lightning density (Ng) of 6.7
flashes/km2/year is taken here (as done in appendix of [37]
for a 35kV tower).
The number of lightning strokes NL (flashes/100 km/year) on
the considered line is calculated using (9). HT and W are the
total tower height and width respectively. Sf is the shielding
factor taken to be 0.5 [37]. The total exposed transmission
segment is 20km (L).
(P) is a cumulative probability distribution function of lightning
stroke in (10) & (11), while If is equal to 28.96kA. Comparing
(10) & (11) indicates a maximum lightning stroke of 110kA on
OHGW over a 12-year period [10]. Equation (10) has been
derived in the appendix of this manuscript. Higher lightning
magnitude, for example 128kA & 137kA occur once in 20 & 25
years respectively for the considered transmission line, have
relatively similar maximum steepness & front time to
110kA lightning strike as estimated from (7) and (8). Thus,
considering the random nature and relatively similar waveform
of large magnitude lightning, 110kA surge can be utilized to
analysis the behavior of OHTL in case of back flashover.
Indirectly Induced-voltage flashovers on the transmission
equipment are not studied here.
B. STEADY-STATE VOLTAGE ACROSS MVDC LINE
TERMINATION
To appropriately add the DC side voltage into the system and
significantly reduce any reflection from the hypothetical
voltage sources, Electromagnetic wave propagation theory is
exploited. Kirchhoff’s laws are valid for characteristic
impedance of any electrical system [37]. For the studied
MVDC link, DC voltage source is connected in parallel
configuration adjacent to one of the converter station
equivalent models in PSCAD/EMTDC. A Bergeron lossless
line with surge impedance 10 times higher than OHTL line is
FIGURE 8. CIGRE lightning stroke with a peak amplitude (I) of 10kA.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
8 VOLUME XX, 2017
FIGURE 9. Characteristic Impedance equivalent model of considered MVDC link.
placed in series with voltage sources to avoid unnecessary
reflection from DC source, as depicted in Fig. 9. Thus,
impact of the emulated DC voltage source is negligible, and
the cumulative characteristic impedance of the converter
station is not impacted significantly. In addition, the
reflection coefficient of the substation busbar & transmission
line tends to reflect 26.06% of substation intruding surge.
While the remainder of it would be refracted into the MMC
substation. Reflection & transmission coefficient between
substation busbar and OHPC are shown in (12) & (13)
respectively [37].
Where, & are reflection and
refraction coefficient for surge transmitting from OHPC to
substation. and are the surge impedances of
transmission OHPC & substation busbar in equation (12) &
(13). Characteristic impedance of components of
transmission link are given in Table. Ⅵ. AC line surge
impedance is added at the ends to eliminate reflections from
model endings.
IV. LIGHTNING STROKE ON TRANSMISSION SYSTEM
The most vulnerable sections of MVDC system from
lightning overvoltage’s are the line transition areas i.e., the
cable section adjacent to the overhead transmission segment
and the substation region [8]. Thus, lightning impact on them
has been studied based on IMSF (10kA) on positive power
conductor and 110kA BF lightning on OHGW.
A. SHIELDING FAILURE AT OHPC ADJACENT TO
CABLE
A scenario is considered where negative polarity lightning
strikes a positive (+ve) polarity pole due to SF at tower (T),
60m away from the riser tower () as shown in Fig. 9. To
study the influence of surge, overvoltage at four other
transmission segments opposite to section is considered
(, , & ). Length of each (voltage measured) tower
section considered is double of its predecessor i.e., is
120m away from T while is 960m.
Lightning strike will initiate a forward voltage surge ()
towards the cable section and reverse voltage surge
towards the transmission line/substation. Initially, total surge
voltage at +ve pole of T (location x) is a sum of & :
will travel across OHPC which is connected to towers via
insulators. will result in negative voltage surge across the
+ve polarity pole insulators. Meanwhile, part of the forward
surge will reflect from riser section and remaining of it
will be refracted into the cable:
The reflected riser surge will be 40% of the initial
value with opposite polarity as estimated from travelling
wave theory [37], equipment surge impedance in Table. Ⅵ
and (15). Remaining 60% percent surge will refract towards
cable entrance . Part of the refracted riser surge
will reflect from the boundary and ultimately reach the
impacted tower section (, depicted as third term in Eq. (16) &
(17). Equation (16) shows, initially, reflected surges retards
the growth of overvoltage () at +ve insulator of tower T.
However, as the steeper part of the lightning surge arrives at
the tower, positive insulator breaks down. Reflection from
the cable end is not accounted for in (16) as it doesn’t
reach instantly due to relatively large cable length. The
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
8 VOLUME XX, 2017
insulation breakdown is represented as a close switch in
PSCAD, governed by equal area criterion model (3).
Overvoltage (n = 1 to 4) at further towers opposite to
riser section (, , & experience similar lightning
overvoltages along with surge retardation due to OHTL
length & towers base. This surge retardation can be
accounted using attenuation constant as stated in (17).
Fig. 10 shows the lightning overvoltage waveform across the
+ve insulator of transmission towers adjacent to UG cable
section due to 10kA shielding failure current at tower T. In
case of T, voltage surge rises initially till 0.2µs after which
reflection from riser section would dampen the voltage surge.
FIGURE 10. Lightning overvoltage at positive pole of adjacent tower
section to the cable.
However, steeper portion of lightning surge (at 5.5µs) results
in breakdown of +ve tower insulator. Fig. 11 depicts the
growth of Insulation disruption coefficient of the considered
tower sections insulators. Once this coefficient goes above
the disruptive area criteria “D”, as recommended for VHD
35-G pin type insulators in Table Ⅳ, the insulators of tower
experience breakdown. In case of further tower i.e., &
surge retardation is imminent as expressed in (17). Due to
relatively short arrival time of reverse travelling surge and
smaller half time () at ’s positive insulator, it doesn’t
experience breakdown.
Once the tower insulator breaks down, surge is transmitted
into tower & ground. Tower surge impedance and
grounding impedance is lower than OHPC this results in
opposite polarity reflected wave to reach the tower top,
retarding the further insulator surge voltage. The simulation
is also done using 200 Ω.m to 800 Ω.m ground resistivity.
Increment in lumped ground resistance (Rg from 64 Ω to 260
Ω) significantly raise the insulation breakdown probability
on farthest towers i.e., & due to lower surge reflections
from the base of adjacent towers as shown in Fig. 12.
B. LIGHTNING SURGE ON OHGW ADJACENT TO
CABLE
This section discusses the backflash-over on OHTL adjacent
to UG cable section. Lightning strikes the ground wire of the
tower “T” for the same configuration of tower as previous
section. Part of the lightning surge traverse along the
impacted tower as well as forward and reverse direction on
OHGW.
FIGURE 11. Prospect of insulation breakdown at positive pole of
considered tower sections based on equal area criteria model.
It can be observed in Fig. 13 that the positive pole
insulator flashover at T occurs before the negative pole under
110kA lightning impact. This is due to fact that tower
insulator with opposite polarity w.r.t tower -ve polarity
lightning overvoltage (), experience largest voltage stress
i.e., . Initially, the +ve insulator
at T do not experience any breakdown as the surge reflected
from tower base ( ) and riser/cable sheath hinders the
growth of flashover until steeper portion of lightning occur at
4.5µs (Fig. 13(a)). After which +ve insulator flashover occur
and part of the surge is injected into the OHPC of +ve pole
which escalate as forward & reverse voltage ().
Part of the forward surge reflects back from the cable/ riser
section to tower section. In case of negative (-ve) pole
insulator, although there is surge attenuation at first due to
positive pole flashover, the large surge half time ( ),
eventually leads to flashover occurring at the -ve insulator of
tower T at 9.8µs. This generates a forward/reverse voltage
surge () on OHPC at –ve pole.
Initially, direct OHGW surge do not cause severe tower
overvoltages at adjacent tower i.e., , & . However,
secondary negative polarity reflected surge from impacted
tower (T) poles and riser/cable section arrive at insulator of
these tower section. Voltage surge (n = 1 to 4) at +ve
insulator for adjacent tower , , & can be
formulated as (18). Similar voltage stress could be observed
on -ve pole insulators as seen in Fig. 13(b). It is seen that
larger half time of CIGRE 110kA lightning strike result in
longer duration of voltage surge at subsequent tower. Thus,
even attenuation constant , due to transmission line length,
do not result in significant reduction of disruptive criterion
coefficient of +ve/-ve insulator (as shown in Fig. 14), except
at tower . Large surge also influence the surge
overvoltages at riser tower . The immediate reflection &
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
VOLUME XX, 2017 9
TABLE Ⅵ
SURGE CHARACTERISTICS OF ±35KV VSC-MMC TRANSMISSION SYSTEM
Equipment
Type
Characteristic Impedance
Symbol
Wave propagation speed
Overhead power cable
147-AL1
538.8 Ω
2.85 × m/s
Overhead ground wire
AWG (6/1) Aluminum conductor
549.5 Ω
2. 85 × m/s
Underground cable core
Single core XLPE cable (± 35)
124.2 Ω
1.75 × m/s
Underground cable sheath
part of cable
23.5 Ω
--
2.02 × m/s
Substation Busbar
--
316 Ω
2.39 × m/s
Hypothetical DC voltage source
--
5000 Ω
--
2.86 × m/s
refraction at cause oscillations at the riser insulators but
larger duration of surge result in insulator flashovers.
The maximum lightning stroke over the transmission line
may vary during its service period. For instance, more
frequent and less prevalent, 31kA & 200kA lightning surge
may strike. Thus, these lightning strikes along with
maximum lightning surge in 20 & 25 years have been
studied. Fig. 15 depicts the peak overvoltages at positive pole
of incident tower, and 18.75km cable entrance and ending for
31kA, 110kA, 128kA, 137kA & 200kA OHGW lightning
stroke. It can be noticed that under the influence of lightning
intrusion, incident tower insulator experience breakdown.
However, for cable section voltage peak are higher for higher
lightning current i.e., for 200kA CIGRE lightning stroke,
cable entrance experiences a voltage impulse above -1000kV
while for 31kA strike, its -178kV. It is expected that increase
in half-time and Sm have resulted in a higher cable entrance
overvoltage.
C. SHIELDING FAILURE LIGHTNING SURGE ACROSS
CABLE
In case of SF, lightning surge partially refract into the
cable section. Positive transmission coefficient generates
negative forward voltage surge at the cable entrance which is
similar in overvoltage polarity as the impacted tower. (19)
represent the forward voltage surge at in terms of
and transmission coefficient at cable junction.
The surge travels and attenuates across the length
of cable section. At cable termination (as depicted in fig.
9) a reflected backward surge is developed. Positive
reflection coefficient from cable and riser junction produces
backward voltage surge with same voltage polarity as
. Super-positioning of these multiple travelling
waves might result in higher initial overvoltage at cable
termination, if surge attenuation (k) of cable isn’t significant
as depicted in (20).
For 18.75km +ve cable pole, 10kA MSF lightning
overvoltages are shown in Fig. 16. The voltages at the cable
entrance, entrance of sections 2, 3 and cable termination are
presented. Propagation time of cable (τ) is large as compared
to lightning surge half time () thus initial surge at cable
entrance subsides before any superposition occur due to
reflections/refractions from cable junctions [6,7]. As depicted
in fig. 16 the first overvoltage surge at cable termination is
higher than at cable entrance due to immediate constructive
interference between forward and reflected waves, resulting
in -205kV peak voltage surge.
D. LIGHTNING VOLTAGE AS A FUNCTION OF CABLE
LENGTH
The impact of cable length on SF lightning voltage has been
studied by considering varying length of the cable from
10km, 5km, 2.5km and 1.25km respectively. It is evident
from Fig. 17 that as the length of the cable decrease (from
Fig. 17 (a) to (d)) there is a significant increase in the cable
termination overvoltage because a shorter cable will not
dampen the surge as much as the bigger segment. This can be
verified using (20) as attenuation constant (k) decrease for
shorter cables. For instance, variation in cable length (from
18.75km to 1.25km) cause peak cable termination
overvoltage reaches up to -300kV breaching the cable
breakdown limit [33]. While peak cable entrance overvoltage
remains same.
With short cables, propagation time (τ) for surge gets smaller
than surge half time () which makes multiple superposition
of reflected waves eminent [6], along any point of cable,
before the first impulse subside i.e., for 10km cable initial
lightning surge reach at 57.1µs as estimated using
FIGURE 12. Equal area criteria coefficient at positive pole of considered
tower sections w.r.t varying tower ground resistance
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
10 VOLUME XX, 2017
FIGURE 13. 110kA Lightning overvoltage’s at tower’s (a) Positive pin type insulator (b) Negative pin type insulator.
FIGURE 14. 110kA lightning flashover measurement using disruptive area criteria at tower’s (a) positive pole insulator (b) Negative pole insulator.
Table. Ⅵ or depicted in fig. 17 (a). With reduced length
subsequent overvoltage maxima & minima become more
prominent. (See fig. 17 (c) & (d)). Cable’s sheath grounding
resistance may vary during its service period. It is observed
that 18.75km cable entrance, under IMSF, endure 6.3%
increase in overvoltage when grounding resistance vary from
10 Ω to 100 Ω. Whereas cable sheath peak overvoltage
drastically rises up to 62% (Table. Ⅶ).
E. EFFECT OF RISER SURGE IMPEDANCE ON
LIGHTNING IMPACT
As described in Ⅱ-G the riser section surge impedance
(Zriser) was taken to be 230Ω for the above considered case.
But it can widely vary between overhead transmission surge
impedance and cable surge impedance owning to the fact that
not much research has been conducted on high-frequency
riser section modelling.
Change in riser characteristic impedance cause variation in
surge refraction & transmission coefficient which impact
and . In fig. 18, it can be observed that for
tower sections other than , lower riser section surge
impedance doesn’t account to significant change in
disruption coefficient for +ve insulator under 10kA lightning
strike. This can be explained based on Ⅳ-A, reflections from
riser section have lesser impact on flashover of OH tower
insulators adjacent to as compared to the steeper portion
of lightning surge. Significant increase in disruption
coefficient at riser section doesn’t cause flashover but may
influence the cable overvoltages.
F. INCORPORATION OF SURGE ARRESTERS AND
SURGE MITIGATION
Usually, for 35kV insulation lightning-withstand-voltage
is about 145 to 170 kV [33]. Recent research papers
recommend a higher insulation level for DC cables as
compared to AC cable of same voltage [2, 14]. PSCAD result
FIGURE 15. Overvoltage prospect of cable segment parts based on
different lightning intrusion.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
VOLUME XX, 2017 11
FIGURE 16. Overvoltage across an 18.75km cable segment due to 10kA
surge on a 60m away tower section.
show that cable might experience breakdown due to SF or
back flashover lightning strikes on OHTL. Thus, an
appropriate surge mitigation scheme is required for the cable
section. A maximum continuous operating voltage (MCOV)
of 39kV is chosen for the surge arrester (1.12 p.u.) [39].
Station low surge arrester (PVI-LP) has been incorporated
for the protection of underground cables [40]. The arresters
lightning Impulse Protection Level (LIPL) is 148kV while
the Switching Impulse Protection level is 105kV. The
protection margin (PM) of the arrester can be calculated as:
PM is calculated to be 15% for the case under consideration.
High-frequency model of a MOSA has been shown in Fig.
19(a) & (b) along with surge arrester's I-V characteristic.
The nonlinear attributes of the surge arrester are
represented as and , shown in Fig. 19(a). In case of a
slow front (switching impulse), low pass filter in fast front
model allows the current to pass through as well as ,
manifesting the character of arrester for low-frequency surge.
However, in case of a fast front surge only is suppressed
resulting high-frequency response of the arrester. Arrester’s
V-I characteristic conversion into and has been
performed according to [40] and parameters of single surge
arrester model with a height of 0.5842m are evaluated as
given in Table. Ⅷ.
On the bases of lightning overvoltage maxima across the
cable section. The most vulnerable part i.e., boundaries of
the cable, are connected with surge arrester. The resultant
forward voltage at the cable entrance as a function
of arrester current and initial riser forward voltage
can be formulated as (done in [6]):
Equation (21) portrays the influence of cable entrance’s
arrester on . govern the surge overvoltage
across the cable. As depicted in Fig. 20 (a), in case of 10kA
shielding failure overvoltage impulse is clipped at each
portion of the 1.25km long cable. Due to low half-time of
IMSF initial overvoltage surge subsides within 50µs. For
OHGW lightning (Fig. 20 (b)), this arrester configuration
limits the cable overvoltage below LIWL of cable. It is seen
that for 1.25m cable surge residual voltage is dictated by
FIGURE 17. Lightning impulse overvoltage at different section for +ve pole underground cable length of (a) 10km (b) 5km (c) 2.5 km (d) 1.25km
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
12 VOLUME XX, 2017
lightning wave-shape/half-time and location of surge.
V. LIGHTNING STROKE ON MMC SUBSTATION
Most likely, the converter & substation are heavily protected
against any direct lightning strokes. However, lightning
overvoltages could traverse through a nearby connected
OHTL. To emulate worst-case scenario, lightning fault on
tower, 60m away from the converter station is considered.
A. LIGHTNING IMPACT ON SUBSTATION SWITCHGEAR
High-frequency MMC substation modelling (as described in
section Ⅱ-C) suggest that AIS busbar might experience
multiple reflection/refraction across it due to difference in
characteristic impedance as compared to OHTL and short
length (25m). In addition, impact of mutual surge impedance
between poles and Earth metallic return have been
considered which implies that lightning surge on one pole
would also impact the remaining busbars.
FIGURE 18. Overvoltage Insulation flashover measurement on positive
pole with varying riser surge impedance under the influence of IMSF.
Fig. 21 shows the lightning overvoltages waveform across
PLD and smoothing reactor in case of 10kA SF current on
nearby +ve pole. The initial oscillating voltage surge at pole
line disconnector remains till 15µs due to small half-time of
IMSF. For smoothing reactor, the oscillating surge is low
because of its parallel stray capacitance in high-frequency
model. Other switchgear equipment like DC circuit breaker,
LD, and SD would experience similar overvoltage due to
short length and symmetrical modelling in PSCAD.
To observe the maximum peak overvoltage, no converter
station arrester has been added. As converter station
switchgear are interconnected with each other by balancing
capacitors & electronic converter structure. The 10kA tower
surge also traverses through negative and ground/earth pole
switchgear. Fig. 22 shows the response of negative and earth
switchgear under the influence of lightning surge at +ve pole.
B. IMPACT OF VOLTAGE SOURCE ON LIGHTNING
SURGE AT MMC CONVERTER
In-order to validate the placement of hypothetical voltage
source, lightning overvoltages at MMC poles with and
TABLE Ⅶ
CABLE AND SHEATH OVERVOLTAGES WITH RESPECT TO VARYING SHEATH
GROUNDING RESISTANCE.
Cable Sheath grounding
resistance (Ω)
Maximum cable
overvoltage (kV)
Maximum sheath
Overvoltage (kV)
10
-190
-32.01
25
-197.05
-39.79
50
-200
47.11
75
-202.55
-50.19
100
-203.96
-51.88
TABLE Ⅷ
PARAMETER OF SURGE ARRESTER MODEL
Components of surge Arrester Model
Magnitude
4.0017 µH
37.973 Ω
0.0068 µH
58.420 Ω
171.17 pF
FIGURE 19. (a) fast front model of surge arrester (b) High frequency V-I characteristic of considered PVI-LP SL arrester for 8/20µs lightning waveshape.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line.
VOLUME XX, 2017 13
FIGURE 20. Lightning overvoltage at 1.25km long positive cables pole with MOSA (on both ends) due to (a) 10kA MSF (b) 110kA OHGW lightning strike.
without it are needed. Ideally, voltage source across the
converter shouldn’t alter the voltage waveform at its poles
except shifting the surge to nominal voltage of +ve/-ve pole.
Fig. 23 shows the lightning voltage waveshape at converter
poles under above discussed scenario for 10kA lightning
strike. Thus, validating the method to incorporate DC voltage
source alongside MMC converter, described in section Ⅲ-B.
FIGURE 21. IMSF lightning resultant voltages at smoothing reactor and
pole line disconnector.
C. LIGHTNING TRANSIENTS ON MMC SUBSTATION
DUE TO BF
Low probability OHGW lightning impact is critical to
measure the maximum overvoltage that could ever be
endured by MMC converter station. 110kA lightning surge is
injected into the OHGW adjacent to the substation. The peak
overvoltages experience by different converter station
sections are illustrated in Fig. 24.
D. LIGHTNING SURGE REDUCTION USING ARRESTER
AT CONVERTER SWITCHGEAR
It is quite clear that substation peripherals could be damaged
due to high magnitude lightning strike or being in close
contact to IMSF. To mitigate the risk of insulation damage
surge arresters are installed at the switchgear entrance
adjacent to tower model in PSCAD along with other arresters
(i.e. DR) specified in Fig. 5. The resulting voltage waveform
at +ve pole disconnector could be noticed in Fig. 25. The
peak overvoltage at other converter station portions has been
depicted in table. Ⅸ. It is tabulated that earthing switchgear
still have surge overvoltage above lightning impulse
withstand level (170kV) due to direct impact from BF.
FIGURE 22. Negative and Earthing switchgear overvoltage under 10kA
lightning surge on nearby positive pole of tower section.
VI. GENERALIZED EVALUATION
To assess the overvoltage surge along an MVDC
transmission line, extensive time domain simulation &
travelling wave theory based numerical analysis have been
carried out and broad range of parameters have been
estimated. Following are the assessments regrading main
characteristics of MVDC system’s lightning behavior:
SF lightning strike might superimpose a -ve polarity
overvoltage surge on positive pole of the tower adjacent
to riser/cable section. Flashovers occur at the impacted
FIGURE 23. 10kA Lightning Surge response of the converter station
poles with and without the hypothetical voltage source.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line
14 VOLUME XX, 2017
FIGURE 24. 110kA lightning overvoltage maxima at different locations of
the converter station without surge arresters.
& adjoining towers’ (except the riser tower) +ve
insulators, regardless of opposite polarity surge
reflections from riser section & underground cables.
BF on such tower results in insulator breakdown at each
of its pole insulators. However, insulator with higher
voltage stresses experience flashover first i.e., +ve
insulator at the impacted tower experience flashover first
due to higher voltage stress from the -ve polarity
lightning surge. TABLE Ⅸ
LIGHTNING PEAK OVERVOLTAGE AFTER ADDING SURGE ARRESTERS AT
EACH POLE OF THE SUBSTATION.
Components of MMC substation
Lightning
magnitude (kA)
Peak Voltage
(kV)
Positive pole switchgear
10
-80.259
Negative pole switchgear
10
-52.272
Earthing switchgear
10
-70.785
Positive MMC converter terminal
110
-43.54
Negative MMC converter terminal
110
-63.98
Negative pole switchgear
110
-133.94
Earthing switchgear
110
-186.64
FIGURE 25. 110kA lightning performance of positive line disconnector at
the substation.
In an event of BF on the tower adjacent to riser section,
lightning surge on OHGW do not cause immediate back
flashover on adjoining towers. This is because part of
the surge is retarded by reflection from tower base.
However, steeper portion of surge from impacted
tower’s poles results in flashover of adjoining tower’s
insulator.
In case of shielding failure and BF, immediate surge
reflections from cable/riser section suppress the
flashover of riser pole insulators. However, due to large
surge half-time (hf) of BF lightning, riser tower poles
ultimately experience flashover.
Likeliness of +ve pole insulators at farthest tower
increase with higher footing resistance due to SF on
tower adjacent to the cable.
A previous study on mixed HVDC transmission line had
shown that riser connected tower is highly resistant to
insulator flashover in case of SF/BF lightning incident
on it [8]. However, it has been identified in this research
that for BF on mixed MVDC transmission line, riser
tower poles are prone to insulator flashover. It is
expected that installing surge arresters on riser/adjacent
tower poles could improve its performance against back
flashover lightning.
Higher riser section characteristic impedance increases
the likelihood of insulator disruption at riser tower.
Without surge arresters at cable joint, it experiences
initial surge overvoltage above its Lightning-withstand-
level due to shielding failure current at adjacent tower
section. Shorter cables tend to have higher terminal
overvoltage as well as secondary maxima/minima due to
reduction in surge retardation factor and lower
propagation time.
Surge arresters at cable terminal clips the lightning
overvoltage across the cable. The surge voltage
waveform at cable entrance is dependent on surge
arrester current , lightning waveform parameters (Sm
& ) & transmission/reflection coefficients of cable.
Lightning surge on MMC substation due to SF on
nearby tower’s +ve pole results in oscillating
overvoltage at impacted busbar pole. Although,
overvoltages are observed at -ve pole and earth pole line
switchgear but are less severe as compared to the
impacted pole switchgear.
A comprehensive method has been proposed in Section
Ⅲ-B to introduce steady-state system voltage, alongside
the MMC converter, in PSCAD simulations for the
lightning study of short transmission lines. The validity
of this method is demonstrated in Ⅴ-B by comparing the
voltage waveform at the converter poles with and
without the hypothetical voltage source. The results
show that adding the voltage source shifts the voltage at
the converter terminal to the nominal voltage without
altering its waveshape.
Substation without surge arrester show extremely high
overvoltage on all poles because of BF on nearby tower.
The surge overvoltages are mitigated by surge arresters’
configuration across the substation. However, lightning
overvoltages still exceed the LIWL at the earth pole
switchgear.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line
14 VOLUME XX, 2017
Finally, it is emphasized that general lightning impulse levels
w.r.t rated system voltage doesn’t seem to be a beneficial
measure. As they are dictated by project specific parameters
i.e. tower structure, grounding condition or cable etc.
Additionally, BF resultant overvoltages are estimated for
110kA lightning current amplitude. However, more severe
lightning might occur at tower adjacent to cable depending
on project specific ground stoke densities, stroke
probabilistic nature [36, 37]. Nevertheless, it is deduced that
OHGW do not prevent insulator breakdown on impacted
towers sections locally in cases of SF/BF. Therefore, in
sensitive transitioning regions of MVDC systems where
permanent current faults due to lightning strike are need to be
avoided, such as OHTL connecting to cables adjacent to a
substation, a few adjoining towers should be equipped with
both OHGW and tower surge arresters [10].
VII. CONCLUSION
As worldwide number of MVDC projects realized with
“Mixed” transmission structure rises continuously, a
profound understanding of lightning stresses affecting the
cable-tower or tower-converter station conjunction are of
major importance. This contribution determines the absolute
maximum lightning impulse voltage waveforms along the
tower section & adjoining underground cable which ranges
from 1.25km to 18.75km. Particularly with regards to the
lightning surge reflection/refractions from overhead power
line, cable and riser section, thorough electromagnetic
transient studies have been carried out for backflash over and
shielding failure lightning. In addition to that overvoltages at
the converter station peripherals have been investigated.
Future studies must focus on devising strategies to suppress
tower insulator breakdown at critical locations of tower-cable
or substation junctions to prevent lightning surge from
resulting in current faults. Future discussions need to clarify
whether standard lightning impulse test practice should be
extended to include superimposed steady-state DC voltages
for tower insulators and cables, considering different DC
voltage levels. The results obtained within this paper are
valuable for insulation coordination of mixed MMC-MVDC
transmission system.
APPENDIX
A. ESTIMATION OF MAXIMUM BF OVER OHTL FOR
RANGE OF YEARS
Maximum lightning magnitude across the overhead
transmission line (OHTL) for a certain period can be
evaluated using its back flashover rate (BFR). Highest BFR
(flashes/100km/year) for a certain lightning magnitude I, at
overhead ground wire directly above tower, is a product of
its cumulative probability and average number of
flashes on the OHTL per 100km per year () [22].
By definition, BFR of a single flash of certain lightning
current magnitude (I) over a period of Y years on air
exposed transmission line of length L (km) can be
expressed as:
Comparing & result in the equation
Thus, 20km air exposed OHTL experiences a 110kA
magnitude lightning strike once in 11.92 or 12 years due to
an of 39.624 flash/100km/year.
ACKNOWLEDGMENT
The authors express their gratitude to the BK21- 4 project,
funded by the Ministry of Education (MOE, Korea) and
National Research Foundation of Korea (NRF).
REFERENCES
[1] S. Coffey, V. Timmers, R. Li, G. Wu, and A. Egea-Alvarez, “Review
of MVDC applications, technologies and future prospects,” Energies,
vol. 14, no. 24, p. 8294, Dec. 2021. doi: 10.3390/en14248294
[2] J. Yu, K. Smith, M. Urizarbarrena, N. MacLeod, R. Bryans, and A.
Moon, “Initial designs for the ANGLE DC project; converting
existing AC cable and overhead line into DC operation,” in 13th IET
International Conference on AC and DC Power Transmission
(ACDC 2017), 2017, pp. 1–6.doi: 10.1049/cp.2017.0002
[3] CIGRE Study Committee C6, “Medium voltage direct current
(MVDC) grid feasibility study,” France, 2020.
[4] Surawijaya, S., Buchner, A. A., & Schichler, U. (2020, September).
Investigation of the Possibility to Convert Medium Voltage AC
Overhead Lines (OHL) to DC. In 2020 55th International
Universities Power Engineering Conference (UPEC) (pp. 1-6). IEEE,
doi: 10.1109/UPEC49904.2020.9209885
[5] Filipović-Grčić, B., Franc, B., Uglešić, I., & Betz, H. D. (2016).
Lightning overvoltage protection of combined overhead line and
underground cable distribution network. In CIGRE International
Colloquium on Lightning and Power Systems, Bologna.
[6] M. Goertz, S. Wenig, C. Hirsching, M. Kahl, M. Suriyah, and T.
Leibfried, ‘‘Analysis of extruded HVDC cable systems exposed to
lightning strokes,’’ IEEE Trans. Power Del., vol. 33, no. 6, pp. 3009–
3018, Dec. 2018. doi: 10.1109/TPWRD.2018.2858569
[7] T. Henriksen, B. Gustavsen, G. Balog, and U. Baur, “Maximum
lightning overvoltage along a cable protected by surge arresters,”
IEEE Trans. Power Del., vol. 20, no. 2, pp. 859–866, Apr. 2005. doi:
10.1109/TPWRD.2005.844262
[8] M. Asif, H.-Y. Lee, U. A. Khan, K. Park, and B. W. Lee, ‘‘Analysis
of transient behavior of mixed high voltage DC transmission line
under lightning strikes,’’ IEEE Access, vol. 7, pp. 7194–7205, 2018,
doi: 10.1109/ACCESS.2018.2889828
[9] Florkowski, M., Furgał, J., & Kuniewski, M., “Lightning Impulse
Overvoltage Propagation in HVDC Meshed Grid,” Energies, vol. 14,
no. 11, p. 3047, Dec. 2021. doi: 10.3390/en14113047
[10] Y. Gao et al., “Study on lightning protection scheme of multiterminal
MMC-MVDC distribution system,” High Voltage, vol. 5, no. 5, pp.
605–613, Jul. 2020. doi: 10.1049/hve.2019.0256
[11] W. Z. El-Khatib, J. Holboell, T. W. Rasmussen, “High frequent
modelling of a modular multilevel converter using passive
components”, International Conference on Power Systems Transients
(IPST2013) in Vancouver, Canada July 18-20, 2013.
[12] R. Zhu, N. Lin, V. Dinavahi, and G. Liang, ‘‘An accurate and fast
method for conducted EMI modeling and simulation of MMC-based
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
M. Usman et al.: Lightning study of a Modular Multilevel Converter based short-length mixed MVDC transmission line
14 VOLUME XX, 2017
HVdc converter station,’’ IEEE Trans. Power Electron., vol. 35, no.
5, pp. 4689–4702, May 2020. doi: 10.1109/TPEL.2019.2945931
[13] CIGRE Working Group B2.62, “Compact DC overhead lines,”
France, 2021.
[14] Y. Liu, X. Cao, M. Fu, “The Upgrading Renovation of an Existing
XLPE Cable Circuit by Conversion of AC Line to DC Operation”,
IEEE PWRD, Vol. 32, Issue 3, pp. 1321 – 1328, 2017. doi:
10.1109/TPWRD.2015.2496178
[15] CIGRE Joint Working Group A3/B4.34, “Technical requirements
and specifications of state-of-the-art HVDC switching equipment”
Technical Brochure, no. 683, 2017.
[16] Y. Han, L. Li, H. Chen, and Y. Lu, “Influence of modeling methods
on the calculated lightning surge overvoltages at a UHVDC
converter station due to backflashover,” IEEE Trans. Power Del.,
vol. 27, no. 3, pp. 1090–1095, Jul. 2012. doi:
10.1109/TPWRD.2012.2190306
[17] M. Rioual : “Short and long air gaps (insulator strings and spark
gaps) modelling for lightning studies with EMTP program (EPRI-
DCG version 2.0)”, Research project, final report, March 88
[18] N. Ahmed, L. Ängquist, A. Antonopoulos, L. Harnefors, S. Norrga,
and H. Nee, “Performance of the modular multilevel converter with
redundant submodules,” in Proc. 41st Annu. Conf. IEEE Ind.
Electron. Soc., Nov. 2015, pp. 3922–3927. doi:
10.1109/IECON.2015.7392712
[19] U. N. Gnanarathna, A. M. Gole, and R. P. Jayasinghe, “Efficient
modeling of modular multilevel hvdc converters (mmc) on
electromagnetic transient simulation programs,” IEEE Transactions
on Power Delivery, vol. 26, no. 1, pp. 316–324, 2011. doi:
10.1109/TPWRD.2010.2060737
[20] HiPak, A. B. B., & Module, I. G. B. T. 5SNA 1200G450350.
[21] M. Rioual, “Measurements and computer simulation of fast
transients through indoor and outdoor substations,” IEEE
Transactions on Power Delivery, vol. 5, no. 1, pp. 117-123, 1990.,
5(1), 117-123. doi: 10.1109/61.107263
[22] CIGRE Working Group 33.01, “Guide to Procedures for Estimating
the Lightning Performance of Transmission Lines,” Technical
Brochure, no. 63, 1991.
[23] Sun, Q., Yang, L., Zheng, Z., Han, J., Wang, Y., & Yao, L., “A
comprehensive lightning surge analysis in offshore wind farm,”
Electric Power Systems Research, vol. 211, 108259, Oct. 2022. doi:
10.1016/j.epsr.2022.108259
[24] H. Rao, “Architecture of Nan’ao multi-terminal VSC-HVDC system
and its multi-functional control,” CSEE Journal of Power and Energy
Systems, vol. 1, no. 1, pp. 9–18, Mar. 2015. doi:
10.17775/CSEEJPES.2015.00002
[25] PROMOTioN., “Task 6.8 Developing a roadmap for VARC DC CB
scaling to EHV DC Voltage,” (Marjon Popov,2019), pp. 1- 54.
[26] T. Joseph et al., “Analysis of harmonic transfer through an MVDC
link,” in Proc. 15th IET Int. Conf. AC DC Power Transmiss., 2019,
pp. 1–6. doi: 10.1049/cp.2019.0017
[27] W. A. Chisholm, Y. l. Chow, and K. D. Srivastava, “Lightning surge
response of transmission towers,” IEEE Trans. Power App. Syst.,
vol. PAS102, no. 9, pp. 3232–3242, Sep. 1983. doi:
10.1109/TPAS.1983.318134
[28] L. Grcev, “Modeling of grounding electrodes under lightning
currents,” IEEE Trans. Electromagn. Compat., vol. 51, no. 3, pp.
559–571, Aug. 2009. doi: 10.1109/TEMC.2009.2025771
[29] T. Watanabe, K. Fuku, H. Motoyama, and T. Noda, “The
measurement and analysis of surge characteristics using miniature
model of air insulated substation,” in Proc. Int. Conf. Power Systems
Transients, 2005. doi: 10.3390/en11092492
[30] ÖNORM E 4101, “Elektrische Freileitungen; Stützenisolatoren VHD
und VHD-G”, Austrian Standard, 1976.
[31] A. F. Imece et al., “Modeling guidelines for fast front transients,”
IEEE Trans. Power Del., vol. 11, no. 1, pp. 493–506, Jan. 1996.
[32] R. D. Caldwell and M. Darveniza, “Experimental and analytical
studies of the effect of non-standard waveshapes on the impulse
strength of external insulation,” IEEE Trans. Power App. Syst., vol.
PAS-92, no. 4, pp. 1420– 1428, Jul. 1973.
[33] Insulation Co-ordination Part 1 Definitions, Principles and Rules,
IEC 60071-1, p. 33, 2006-01.
[34] Asif, M., Lee, H. Y., Park, K. H., & Lee, B. W., “Accurate
evaluation of steady-state sheath voltage and current in HVDC cable
using electromagnetic transient simulation,” Energies, vol. 12, no.
21, p. 4161, Oct. 2019. doi: 10.3390/en12214161
[35] Z. G. Datsios, P. N. Mikropoulos, and T. E. Tsovilis, “Effects of
lightning channel equivalent impedance on lightning performance of
overhead transmission lines,” IEEE Trans. Electromagn. Compat.,
vol. 61, no. 3, pp. 623–630, Jun. 2019.
[36] IEEE Guide for improving the Lightning performance of
Transmission Line, IEEE Std. 1243-1997, Dec. 1997.
[37] M. Paolone, R. Farhad, and C. A. Nucci, IEEE Guide for Improving
the Lightning Performance of Electric Power Overhead Distribution
Lines. Piscataway, NJ, USA: IEEE, 2010. no. EPFL-STANDARD-
181604.
[38] “Cables Overvoltage protection – ABB”,
https://search.abb.com/library/, accessed 5 June 2023.
[39] ‘Station Class Surge Arresters IEEE and IEC’,
https://hubbellcdn.com/catalogfull, accessed 5 June 2023.
[40] ‘Metal Oxide Surge Arrester for PSCAD Version 5.0’,
https://www.pscad.com/knowledge-base/article/609, accessed 5 June
2023.
MUHAMMAD USMAN was born in Multan,
Pakistan, on March 30, 1998. He graduated from
University of Engineering and Technology
(UET), Lahore (Pakistan) in 2019 with bachelor’s
degree in Electrical Engineering.
Currently he is enrolled in MS-PhD program at
HVDC Electric Power Lab, Hanyang University,
Ansan, South Korea. His fields of interest include
power electronics, insulation coordination studies
for DC grids, protection system and switchgear
design for HVDC & MVDC system.
Mr. Usman is a Student Member of Korean Institute of Electrical
Engineers (KIEE).
KYU-HOON PARK (Student Member, IEEE)
was born in Seoul, South Korea, on March 20,
1991. He graduated with bachelor’s degree in
Electronic System Engineering from the Hanyang
University (South Korea) in 2016.
Currently he is a combined MS-PhD candidate
at HVDC Electric Power Lab, Hanyang
University, Ansan, South Korea. His special
fields of interest include Power System Analysis
and Protection Equipment for dc grids.
Mr. Park is a Student Member of CIGRE and
Korean Institute of Electrical Engineers.
BANG-WOOK LEE (Senior Member, IEEE)
received the B.S., M.S., and Ph.D. degrees from
the Department of Electrical Engineering,
Hanyang University, Seoul, South Korea, in
1991, 1993, and 1998, respectively. He was a
Senior Research Engineer at LS Industrial
Systems Company, Ltd., South Korea. In 2008,
he joined the Department of Electronic
Engineering, Hanyang University, Ansan, South
Korea, where he is currently a professor. His
research interests include HVDC protection
systems, high voltage insulation, renewable energies, development of
electrical equipment, and transmission line structures for HVDC and
HVAC power systems. He is a member of HVDC Research Committee of
KIEE, Power Cable Experts Committee of the Korean Agency for
Technology and Standards, and CIGRE.
This article has been accepted for publication in IEEE Access. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/ACCESS.2023.3293531
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/
