Energy and Power E ngineering, 2013, 5, 1266-1272
doi:10.4236/epe.2013.54B240 Published Online July 2013 (http://www.scirp.org/journal/epe)
Copyright © 2013 S ciRes. EPE
Analysis of a Large Groundin g System and Subsequent
Field Test Validation Using the Fall of Potential Method
Huan Huang1, Hualin Liu1, Hong Luo1, Hao Du1, Yi Xing1, Yexu Li2,
Farid P. Dawalibi2, Haijun Zhou2, Longhai Fu2
1Guizho u Electri c Power Test & Research Institute, Guiyang, China
2Safe Engin eering Services & technologies ltd, Blvd. Des Ois eaux, Laval, Québec, Canada
Received April, 2013
ABSTRACT
This paper examines various aspects of the design process and subsequent field test measurements of a large and com-
plex s ub st at io n gro u nd in g s yst e m. T he st ud y a nd mea sur e ments show that soil layering and lead interference can have a
significant impact on the appropriate test location that yields the exact substation ground impedance. Applying a spe-
cific percentage rule such as the 61.8% rule for uniform soils to obtain the true ground impedance may lead to unac-
cep ta bl e e rr or s fo r l ar ge gr o un di ng s ys te ms. T hi s p o ses s i gni fic a nt p r ob l ems whe n a ttempting to va li da te a d e sig n b ase d
on raw test data that are interpreted using approximate methods to evaluate substation ground impedance, and determine
ground potential rise (GPR), touch and step voltages. Advanced measurement methodologies and modern software
packages were used to obtain and effectively analyze fall of potential test data, compute fault current distribution, and
evalua te touch and step voltages for this lar ge substation. Fault current distribution between the grounding system and
other metallic paths were computed to determine the portion of fault current discharged in the grounding system. The
performance of the grounding system, including its GPR and touch and step voltages, has been accura tely computed and
measured, taking into account the impedance of the steel material used of the ground conductors and circulating cur-
rents withi n the sub s tat io n gr ound i n g syst e m.
Keywords: Ground Resista nce; Fal l-of-Potential; Ground Impedance Measure ment; Ground Potential Rise; Ground
Potential Difference; Touch Voltage; Step Voltage; Steel Conductors
1. Introduction
Appropriate power system grounding is important for
maintaining reliable operation of electric power systems,
protecting equipment, and insuring the safety of public
and personnel. A grounding system must be properly
designed and its performance needs to be evaluated. Im-
proper or inaccurate analysis can lead to significant ex-
penses due directly to unnecessary over design or as a
result of subsequent corrective measures caused by fail-
ures of the inadequate design. Most electrical engineers
understand the importance of grounding system to dis-
charge safely phase-to-ground faults into the sur round in g
soil.
Unfortunately, the complex and non-homogeneous
nature of the soil, the intricate three-dimensional shape of
the grounding system and topology of the entire power
system network result in a very difficult ta sk that requires
appropriate specialized software packages and skilled
professionals with adequate expertise in this field in
order to account for the numero us factors that have t o be
considered during the design process and subsequent
field measure ment validation task.
It is often necessar y to measure the ground impedance
of a grounding system in order to validate a grounding
analysis. The basic technique which is almost universally
used for the measurement of grounding system
imped ance is known a s the Fal l-of-Potential method. T he
Fall-of-Potential method introduces two auxiliary
electrodes, called return electrode and potential probe.
When the return electrode is placed at a finite distance
from the grounding system and the potential probe is
driven into the earth at a specific location(the so-called
“exact potential probe location”) then an accurate
measurement of the ground imp edance is obtained.
For uniform soils and large distances between the
grounding system and the return electrode, it is well
known that t he exact poten tial prob e location follo ws the
61.8% rule, i.e., the exact locatio n for the potential prob e
is r = 0.618D, where r and D are the distances from the
center of the gro undi ng s yste m to the p ot entia l pr obe and
to the current electrode, respectively. It is important to
understand that the 61.8% rule is based on the

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1267
assumption that the soil is uniform and that the
grounding electrode is small or hemispherical and that
the potential probe and current electrode are also
hemispherical or small.
For two-layer soils, a classical paper [3] has shown
that the exact locations can vary from values close to
50% to values exceeding 90% depending on the nature of
the t wo -l ayer s o il struct ure an d thick ness of the top la yer.
This finding was included in Guide IEEE 81 -1983 [15]
and was mysteriously removed in the 2001 edition before
being reintroduced in the 2013 edition! In other words,
the exact location of the potential probe is well defined
for some ideal cases, such as hemispherical or small
grounding electrodes buried in uniform or layered soils
[1-6] but must be evaluated adequately when the
separation distances are not large enough. In such cases,
a value read at 61.8% may lead to significant errors on
the measured ground impedance. The exact potential
probe position must therefore be determined each time,
using appropriate computer simulations.
Many grounding systems in China and several other
countries are made of steel, which has higher
permeability and lower conductivity than copper [7-11].
This raises some unique issues, particularly if the
substation size is lar ge and the soil resisti vity is lo w. In a
conventional grounding analysis approach, a grounding
system is generally assumed as an equipotent structure.
This would be inaccurate for most cases where steel
grounding systems are used. In fact, the ground
impedance of the grounding system has a significant
inductive component, which is not take n into account b y
classical groundin g analysis metho ds.
The analysis of the grounding system of an existing
500 kV large substation is summarized in this paper. The
substation includes 500 kV, 220 kV and 35 kV
switchyards. T wo500 kV and seve n 220 kV transmission
lines enter the substation.
Figure 1 is a plan view of the substation grounding
system. The ground conductors are buried at a depth of
0.8 m and are made of L60*6 mm steel conductors. A
number of ground rods are installed at various locations
of the gr id . They are 2 .5 m long and are made of L60*3.5
mm steel conductors.
Figure 2 is a plan view of the substation grounding
system and the electrical network connected to it. Figure
2 also provides the soil resist ivity measure ment locations:
one is inside the substation while another one is outside
the substatio n. Figure 3 represents the multiphase circuit
for a single-line-to-gro und fault in the 2 20 kV sub statio n
yard. It shows the equivalent circuit of the computer
model used.
Figure 4 shows a typical cross section of all the
transmission line towers modeled. A tower resistance of
15 or 20 ohms was used depending on the type of tower
structure grou nds.
Figure 1 . Plan view of the g r o unding sy stem.
Figure 2. Plan view of the grounding system and the net-
w ork connected to it.
Figure 3. Simplified circuit model for fault current split
calculations.

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Figure 4 . Typical cross section of the transim msiosn lines.
To evaluate the grounding performance of the large
syste m ground net work, the follo wing steps were carried
out: 1) Soil resistivity and grid impedance measurements
and interpretation; 2) Fault current split calculations; 3)
Groundin g system performance analysis.
This pap er is not intended to repor t detailed grounding
design issues and results but rather to highlight the vari-
ous challenges encountered and their ramifications, if
these challenges are ignored or if simplifying assump-
tions are made in place of a detailed grounding analysis.
The results presented in this paper provide useful in-
sight and information for accurately measuring the
ground impedance of large grounding systems, for inter-
pre ting the mea sure ment s and for evalua ti ng safe ty in t he
station. The analysis and the discussions can be used as a
reference guide to st udy lar ge grounding sys t ems.
2. Resistivity Measurements and
Interp ret ation
Soil resistivity measurements were made along two
trave r ses a t t he s ub sta t io n si te , usi ng the W enne r fo ur -pin
method. Measurements along the short traverse inside the
substation were carried out in order to obtain shallow
depth resisti vities at the pro ject site. Measure ments alon g
one long traverse outside the substation were carried out
in order to obtain soil resistivitie s at la rger depths.
The measured soil resistivity data were interpreted us-
ing the RESAP computation module of the CDEGS
software package [12]. Based on the principle that short
traverse measurements determine shallow depth soil re-
sistivity and large spacing measurements determine deep
soil resistivity, a three-layer soil model was constructed,
which is representative of the soil structures at the site
and is expected to be conservative for the grounding
analysis. The selec te d soil mo del is shown in Table 1.
3. Grounding Impedance Measurement and
Interp ret ation
To evaluate the performance of a substation grounding
system, the ground impedance of the grounding system
must be obtained either by measurement or by com-
putation with appropriate soil resistivity measurements.
Incorrect ground impedance will lead to incorrect fault
curr ent co mputati on, ther efore affecting the results of the
analysis. Ideally, the ground impedance should be
computed and then validat ed by measurement.
For this practical case study, due to the dense
transmission line area and the expected low impedance
measurement, a high current test unit was used.
According to IEEE 81.2, it is recommended that the
current probe has a minimum length of 6.5 times the
diagonal of the substation. The substation is surrounded
by difficult terrains. Fortunately, a small road was
available along the west side of the substation. Therefore,
the fall of potential test that followed this road was used
to perform the measurements. Figures 5 and 6 illustrate
the test setup which was carried out for multiple
frequencies of 47, 46, 53 and 54 Hz for each test point.
The fall of potential test results are summarized in
Figure 7 (red dots).
The next stage of the analysis was to model the fall of
potential test, using the soil model developed from
measured soil resistivity data (shown in Table 1). The
first challenge was to determine how much of the overall
transmiss ion line gr ounding network is re ally i nflue ncin g
Table 1. Selected soil model.
Layer Resistivity (
Ω
-m) Thickness (m)
Top 320 0.35
Ce ntral 65 2.0
Bottom 420 infinite
Figure 5 . Fall of potenti al test setup (complet e model).

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Figure 6 . Fall of potential test setup(zoomed model).
Figure 7 . Computed a nd measured appearent resista nc e.
the GPR of the substation. After building an accurate
model of 5-10 km shield wires, incl udi ng all to wer foo ting
structures and substation grounding grids corresponding
to the real network in the area, the fall of potential r es ults
were computed. A sketch of the model is shown in
Figure 5. In this model, no current or potential leads
were modeled. The MALZ computation module of
CDGES was used. This module takes into account the
voltage drops along a grounding system and is therefore
capable of modeling large grounding systems with steel
conductors. It means that the developed model accounts
correctly for the conductive coupling, but does not
account for inductive effects between the measuring
leads and other paralleling conductors.
The results of this fall of potential simulation are
sho wn in Figure 7(green curv e). As it can be seen, a sig-
nificant discrepancy exists between the measured curve
and the computed curve. The first common sense reac-
tion may be to conclude that the soil model was incorrect.
However, after modeling the test leads (Figures 5 and 6)
and using the HIFREQ computation module of CDEGS
which accounts for all relevant electromagnetic effects,
i.e., for conductive, inductive as well as capacitive cou-
pling effects, one can notice that the resulting theoretical
fall of potential te st results, co mpared with the measured
resul ts (sho wn in Figure 7 as a blue curve and red dots),
agree with each other very well. Conductive coupling
arises as a result of the proximity of current circ uit return
ground grids and nearb y buried structures, such as tower
footing, connected to the ground grid. Inductive coupling
is a result of the test leads being mutually coupled with
buried structures that are connected to the substation
ground grid. Capacitive coupling is due to the capaci-
tance between buried and above ground conductors).
The computed true grounding impedance at 50 Hz is
0.3066∠13.550 Ω for the complete network and0.6349∠
1.210 Ω for the substation grid alone. It is clear that the
exact potential probe location for the entire network im-
pedance, energized at 50 Hz case, is at 33% of the dis-
tance between the injection point and the return electrode,
whic h is fa r from the s ugge s te d r ule o f 61. 8 %. A value o f
0.51Ω is obtained for the ground impedance of the entire
system if the 61.8% reading is used, an error ofabout
66%. This is because a) the return electrode is not located
far enough from the grounding grid, compared to the
measured grounding system dimension (i.e., the complete
network); b) the soil model is non-uniform; c) the test
leads were close and parallel to each other (~4m).
The above analysis has shown how to use a judicious
ope r ati ng fre q ue nc y to mea s ur e the gro und i ng i mp ed a nce
of a substation grid that is connected to other grounding
electrodes through shield wires and interpret the data
correctly. This is an important issue when dealing with
large substa t ion grounding systems.
4. Fault Current Split Calculations
The objective of the fault current split calculations is to
obtain the earth current (current discharged by the
grounding system to earth). Under most conditions, the
total fault current doesn’t discharge entirely in the
substation grounding system. Part of the fault current,
which does not contribute to the GPR of the grid, will
return to remote source terminals and to transformer
neutra l s t hr o ug h shi e ld wir es, neut ra l wire s o r co nd uct or s
of the grid.
It is well known that the GPR and the touch and step
voltages associated with the grounding network are
directl y proportional to the magnitude of the fault current
component discharged directly into the soil by the

H. HUANG ET AL.
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ground in g net work. It i s therefore important to determine
how much of the fault current returns to remote sources
via overhead ground wires and neutral wires of the
transmission lines and distribution lines connected to the
substation.
Computer simulations have been performed using the
Right-Of-Way software package described in [12] based
on the circuit model sho wn in Figure 3 and on the com-
puted ground impedance of the substation in the soil
shown in Table 1. A circuit model representing the sce-
nario of a 220 kV single-line-to-ground fault at the power
plant is sho wn in F igure 3. Figure 3 al so sho ws the fa ult
current contributions from all sources; Figure 8 shows
the distrib ution of the fault curre nt along the tr a ns mission
line overhead ground wires for a 220 kV single-phase-
to-ground fault. Table 2 shows the results of fault curre nt
distribution calculation. Note that the local source con-
tribution from the 500 kV /220 kV step-up transformers
in the substation i s estimated to be about 8 kA, as shown
in Table 2. This current was not modeled in the circuit
current split calculations because it circulates [13] be-
tween the fault location and the step-up transformers via
neutra l and gr ound cond uctor s. Howeve r, this cir culati ng
curr ent was inc lud ed i n the gr oundi ng s yste m mod el a s it
should.
Figure 8. Computed fault current in the power line over-
head ground wires.
Table 2. F a ult c ur rent spli t calculation results.
Remote contribution 9.8
∠
-97.1° kA
Local contrib utio n (circulating current) 8.0
∠
-90.0°kA
Total fault current 17.77
∠
-93.9°kA
Current returni ng via OHGW 8.14
∠
95.5° kA
Earth current discharged in grid 1.78
∠
105° kA
5. Grounding System Performance Analysis
GPR, touch and step voltages are important quantities
when a substation is assessed. The calculation of GPR,
touch a nd step vo ltages was car ried out using the M ALZ
computatio n module[12], which takes into account
attenuation or voltage drop along conductors in a grounding
system, avoiding therefore the incorrect assumption that
a groundi ng grid is equipot ent ial.
In this study case, we have a 500 kV substation at
different voltage levels, i.e., 500 kV, 220 kV and 35 kV,
fed by several power source terminals at 500 kV and 220
kV. Whe n a sin gle-phase-to-ground fault occurs on a 220
kV bus, the 500 kV side will supply fault currents
thro ugh t he 50 0 kV transfo r mers; par t of the fault c urre nt
will return to the transformer neutral through the grid
conductors (circulating current between the fault location
and the trans former neutral po int).
Similarly, the fault current is injected at the fault
location and a portion of it returns to the remote source
through the shield wire con nected to the grid (circulatin g
between the fault location and the shield wire connecting
point). Because of the low impedance path provided by
the ground conduct ors , onl y a ne gligi ble s mal l a mount o f
this c ur re nt le a ks o ut from the gr o und c o nd ucto r s to e ar th.
As a result, cir c ulating curre nts do not affect s ignificantly
the average grid GPR but may distort its shape significantly.
However, for a large substation, the distance between the
fault location and the transformer neutral point or the
shield wire connecting connection point can be quite
large. As a result, high potential differences, due to this
large circulating current, may exist within the grounding
grid. In other words, t he circul ating curr ent co ntribute s to
the ground potential differences (GPD) between various
locations o f the grid and results in hi gher touch and step
volta ge s, es pe ci al l y for a lar ge gro und i ng s yst e m a nd low
soil resistivity soil environments.
Obviously, ignoring transformer and shield or neutral
wirescirculating currents in a groundi ng syst em study can
lead to inaccurate designs leading to unsafe situations.
A co mputer mod el was buil t f or t he gro undi ng sys te m.
Figure 9 shows the model considering all current sources,
includin g circul ating currents.
The maximum acceptable touch and step voltages are
indicated in Table 3 that are calculated based on the fol-
lowing criter ia :
Method: IEEE Standard 80-2000 [14]
Body weight: 50 kg
Body resistance: 1000 
X/R ratio: 20
Fault duration: 0.12 sec
All accessible areas inside and outside the substation,
for all possible soil surface covering materials: native
soil, crushed rock or asphalt (even very low resistivity
materials), had acceptable touch and step voltages under

H. HUANG ET AL.
Copyright © 2013 S ciRes. EPE
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all fault scenarios, i.e., 500 kV or 220 kV faults, assum-
ing computed and measured tower resistances, and at
different fault locations.
Figure 10 shows typical ground potential rises (GPRs)
along the grid conductors for a 220 kV fault. Figure 11
and Figure 12 show typical touch and step voltages in-
side and outside the substation, respectively. Figure 13
provides example of soil pote ntials.
6. Conclusions
The performance of a large substation grounding system
has been analyzed using modern techniques. A non-
uniform soil model has been derived based on soil resistivity
measurements, and it has been applied throughout the
study.
The paper shows that by following the IEEE 81.2
recommended methodologies for fall of potential testing,
significant errors are introduced for a large grounding
system connected to an extended network, due to
conductive and inductive effects. It was shown that
matching the fall of potential test data with a detailed
computer model using an appropriate software package
can significantly change the computed ground potential
rise , touch and s t ep voltages in a s ubstation.
Fiugre 9. Ground network model for evaluating safety at
the substa tion.
Table 3. Safety Limits.
Surface
Maximum Acceptable
Soil
Resistivity
(-m)
Touch
Voltage (V)
Step
Voltage (V)
Very Low 0 272 272
Native
320
407
815
Crushed Rock
2000
973
3078
Wet Concrete
30
282
323
Asphalt 10,000 4518 17257
Figure 10. Conductor GPR (Ground Potential Ri s e) .
Figure 11. Touch voltages at the su bs tation.
Figure 12. Step voltages inside and outside t he substation.

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Copyright © 2013 S ciRes. EPE
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Figure 13. Soil Potentials inside and outside the s ubstation.
A complete circuit model of the overhead transmission
line network has been built in order to determine the
current distribution during a single-phase-to-gr ound fa ult.
Therefore, current injected into the soil through the grid
(which contributes to the GPR, touch and step voltages)
was obtained. Due to the large size of the grounding
system and to the fact that the grid is made of steel
ground conductors, the conventional approach used in
grounding analysis (equipotential grounding system) can
lead to wrong results. Therefore, adequate methods
taking into account voltage drops along the grid
conductors and circulating currents within the substation
must and have been used to compute the grid GPR, touch
and st ep voltages.
The procedures presented in this paper can be used as
a guide when carrying out grounding analysis of a large
po wer substatio n.
7. Acknowled gements
The authors wish to thank Mr. J. L. Chagas of SES for
his review and comments on the paper man uscrip t.
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