1
AC-corrosion on cathodically protected pipelines: A discussion of the
involved processes and their consequences on mitigation measures
Markus BÜCHLER and David JOOS
SGK Swiss Society for Corrosion Protection,
Switzerland, [email protected]
Abstract
The first damages on cathodically protected pipelines occurred in 1988. The subsequent research of
the past 28 years has resulted in a conclusive understanding of the involved processes. The most recent
results are completing this picture. This updated model allows for the design of mitigation measures
that go beyond the decrease of the induced ac-voltage. The depth of the ac-corrosion attack and the
associated critical coating defect size in combination with the soil resistivity were identified as key
parameters with respect to damages on pipelines. The corresponding model concepts are discussed and
their validation presented. Moreover, the consequences on the threshold values in ISO 18086 as well
as the effect of time variation of ac-and dc-interference are addressed.
Keywords
ac-corrosion, corrosion rate, cathodic protection, model
Introduction
Within two extensive DVGW research projects [1, 2] the interference values specified in EN
15280 and ISO 18086 were confirmed by laboratory and field investigations. Unfortunately
satisfying the current density requirements at increased ac-interference levels, high protection
current demand of the pipeline, heterogeneous soil resistivity, and dc-interference conditions
was found to be challenging in the practical application. This is mainly a result of the required
less negative on-potentials (Eon) that do not leave room for required adjustment of the
rectifiers. This inherently bears the risk that the shifting of the Eon to less negative values
results in an important decrease of the ac-corrosion risk, but increases the risk of not meeting
the requirements of EN 12954 and ISO 15589-1 with respect to cathodic protection (CP)
effectiveness. Hence, controlling ac-corrosion can readily result in corrosion due to
insufficient CP.
Within the DVGW research projects corrosion rates higher than 100 mm/year were observed
on electrical resistance coupons (ER-coupons) demonstrating that the control of ac-corrosion
has highest priority. However, these high corrosion rates were never present for longer
periods of time on these pipelines, since they were operated for several decades under these
conditions and would have shown leaks. This discrepancy between coupon data and operation
experience was attributed to the specific geometry of the polymer adjacent to the ER-coupons
or temperature effects caused by the discharge of high current densities. However, further
investigation made clear that neither effect can explain the relevant difference in corrosion
rate on the coupon and the pipeline.
Only more recent investigations based on the numerical description of the effects taking place
under cathodic protection, provided a conclusive explanation for this discrepancy [3-5]. These
model calculations result in two key conclusions:
It is impossible to prevent ac-corrosion at smallest coating defects
ac-corrosion will stop at a certain depth, provided the wall thickness is sufficiently large
2
Based on these observations it was concluded that the ER-coupons with a thickness in the
range of one millimeter correctly describe the ac-corrosion rate in the first stages of the
corrosion process. The decrease of the corrosion rate caused by the increase of the metallic
surface as a result of the propagation of the corrosion process into depth was not possible to
detect due to their small thickness. Hence, the extrapolated corrosion rates on thin coupons
are much too high and may not be used for integrity assessment of pipelines with larger wall
thicknesses.
Since ac-corrosion on small coating defects will even occur at very low induced ac-voltages
(Uac) the occurrence of ac-corrosion cannot be avoided. Alternatively to the corrosion rate, the
proposed model identified the acceptable corrosion depth as the key controlling parameter.
Based on this concept an increased tolerated corrosion depth is expected to allow for more
negative Eon and higher Uac without leaks occurring on the pipeline.
This model concept was validated within a third DVGW (German Technical and Scientific
Association for Gas and Water) research project [6]. The present paper summarizes the
obtained key results. A summary of all used abbreviations is given in the annex.
The model
Since the occurrence of the first ac-corrosion damages in the year 1988 on cathodically
protected pipelines [7, 8] the phenomenon was investigated in detail. Soon the ac-current
density was identified as a critical parameter [9-11]. Additionally it was found that the
protection current density has an important effect on the corrosion process [12-15]. Various
investigations were performed in order to obtain a profound understanding of the relevant
parameters [16-18]. The resulting model concept is capable of explaining all empirical
observations. Within extensive field investigations it was possible to confirm the threshold
values determined in laboratory investigations under realistic operation and interference
conditions [1, 2].
Based on these data the ac-corrosion rate is small, if the average ac-current density (Jac) is
smaller than 30 A/m2 or the average protection current density (Jdc) is smaller than 1 A/m
2. It
was found that the latter is only possible if the average Eon is more positive than -1.2 VCSE and
the average Uac is smaller than 15 V. Additionally, the IR-free potential had to be more
negative than the protection criteria in EN 12954 or ISO 15589-1. Based on the model
concept and the experimental data it was possible to demonstrate that ac-corrosion could even
be prevented at high Jdc, provided Jdc is at least one third of Jac. Based on the currently
available information a profound understanding of the involved processes and the required
threshold values is possible as described in detail in [18].
Relevant parameters
The model for describing the process of ac-corrosion was already discussed in detail [18, 19]
and will not be repeated. However the key controlling factors relevant for the numerical
calculation will be presented. The following observations are characteristic for ac-corrosion
on cathodically protected pipelines [9]:
The pH at the steel surface is significantly increased
Compact corrosion products are formed that mainly consist of Magnetite and Goethite
The corrosion products form directly at the steel surface and result in a separation of the coating from the steel surface
No soluble corrosion products are observed It is known that the cathodic reduction of the passive film formed on the steel surface in
alkaline environment results in the formation of an iron oxide or hydroxide [20].Various
authors have reported the change of the oxidation state of this rust layer under the influence of
an electrical current [21-25] on the passive steel surface. This rust layer plays a crucial role in
3
the process of ac-corrosion at less negative Eon in the range of -1.2 VCSE, since it consumes
the electrical charge during the anodic and cathodic half wave of the ac-current by means of
the redox system Fe(II)/Fe(III). Hence, even significantly increased Jac can pass through the
passive steel surface without resulting in corrosion. A corrosion process will only take place
at increased Jdc (larger than 1 A/m2), which result in the cathodic polarization into the
immunity domain and the cathodic dissolution of the passive film. This effect is illustrated in
Fig. 1 based on the data of Bette [26] for the fast measured IR-free potential of a coupon at
two different Jdc.
The IR-free potential
The behavior in Fig. 1 is of key relevance for cathodic protection and the associated
protection criteria. Despite of significantly increased Jdc the IR-free potential is temporarily
anodic of the protection criterion of -0.85 VCSE. Based on the concurrently recorded corrosion
rate data only a metal loss is observed when a temporary polarization cathodic of -1.2 VCSE
occurs. This effect is fully in line with the polarization into immunity and cathodic dissolution
of the passive film [18]. Special attention has to be paid to the time dependence of IR-free
potential, which shows a variation in the range of 0.4 V as a function of the polarity of the
current density (J). Based on the data in Fig. 1 this time dependence can be assessed with a
data acquisition rate of at least 1 Hz. Slower measuring rates of less than 10 Hz, as they are
usually applied in CP for assessing the IR-free potential do not show this time dependence.
Hence, the question regarding the physical significance of the various IR-free potentials in
Fig. 1 rises. In this context only the most relevant influencing factors and the corresponding
conclusions will be presented. A more detailed discussion of the involved processes is given
in [27].
Fig. 1: IR-free potentials determined on an ER coupon as a function of the current
density (J) under ac-interference of 16.7 Hz (cathodic currents with a positive sign).
Black: No corrosion at Jdc 1 A/m2, Jac 128 A/m
2; Red: Corrosion at Jdc 11 A/m
2, Jac 309
A/m2 according to [26].
Based on Fig. 1 it is not possible to determine a single value for the IR-free potential. It is,
however, possible to determine an average of all the recorded values, which in a first
approach corresponds to the classical slowly measured (about 100 ms after interrupting Jdc)
IR-free potential. This average value will in the further context be described as EIR-free.
4
Additionally, in Fig. 1 also the most negative potential excursion limited by hydrogen
evolution can be determined, which according to the model concept controls the ac-corrosion
process. This parameter is in the following context described as EH.
In absence of ac-interference equation (1) applies:
HfreeIR EE (1)
In presence of ac-interference equation (2) applies with the contribution of the so called
faradic rectification ΔEF [28].
FHfreeIR EEE (2)
This consideration represents a rough simplification of the processes taking place at the steel
surface under exclusion of all time dependent contributions. However, it provides a physical
description for the empirically observed effects caused by the rectification of Jac. The size and
the sign of ΔEF are dependent on the ratio of the Tafel slopes of the anodic and cathodic
partial reactions. It is characteristic for passive systems that ΔEF has a positive sign. This
effect was used for so called "wet rectifiers" (or electrolytic rectifiers). The key requirement
for the electrodes was their passivity (e.g. [29-31]).
The model
The Jdc, which passes through a coating defect with a metallic surface of A, is a result of the
difference between Eon and EIR-free as well as the spread resistance R according to equation
(3). For Eon more positive than -1.2 VCSE it was demonstrated that Jdc can go towards zero
[18]. Based on this concept it is possible to limit ac-corrosion through the control of Eon even
at increased ac-interference.
AR
EEJ
onfreeIR
dc
(3)
The Jac passing through the metal surface causes a shift of the EIR-free in positive direction
according to equation (2). For the determination of Jdc the EIR-free (average IR-free potential) is
relevant, which is a result of EH and ΔEF.
The evaluation of the literature data [19, 32] with respect to Jac and the resulting anodic shift
of EIR-free under assumption of a linear behavior allows describing ΔEF with a factor f
according to equation (4) [3-5].
AR
UfE acF
(4)
fEEARJARU onHdcac (5)
The combination of equations (3) and (4) results in equation (5), which is a description of the
Eon and the allowable Uac as a function of the critical Jdc in the case of cathodic protection at a
less negative Eon.
The further consideration of the thermodynamic [33] and kinetic [25, 34] parameters with a
mathematical description of the decrease of the spread resistance caused by the increase of the
pH-value at the steel surface [19] caused by Jdc allows for a more detailed description of the
corrosion process under ac-interference based on equation (5) [3-5].
5
Conclusions
The present consideration allows for explaining the relevant discrepancy between the actual
damages on pipelines and the high corrosion rate on coupons. Based on equation (5) the
acceptable Uac goes towards zero for decreasing defect sizes. In contrast, an increase of the
metallic surface A in the coating defect will increase the acceptable Uac. If the soil resistivity,
the average Eon and Uac, the original coating defect size and the allowable corrosion depth are
known, an assessment of the acceptable interference conditions is possible. The key
conclusions for ac-corrosion at less negative Eon are therefore as follows:
At small coating defects ac-corrosion cannot be prevented
ac-corrosion should stop at a certain depth
This depth is larger on large coating defects than on small coating defects These qualitative conclusions are in good agreement with the empirical observation of the last
decades. For the practical application of these parameters, however, it is of importance to
quantify the individual parameters.
Validation of the electrical aspects
The individual parameters in equation (5) were investigated for the reliable calculation of the
corrosion behavior and especially the maximum corrosion depth. This will be discussed in the
following.
The calculation
The calculation of EH was performed under the assumption of anaerobic conditions and
increased Jdc of more than 1 A/m
2 according to equation (6) for activation controlled hydrogen
evolution. Cathodic Jdc (when EH is negative of E0) exhibit a positive sign.
kHdc KEEJJ 00 exp (6)
Additionally E0 [VCSE] is calculated according to equation (7).
pHE 0591.032.00 (7)
Based on equation (6) and (7) EH can be determined at a known pH-value for every Jdc.
The reduction of water results in an increase of the concentration of OH- at the steel surface.
The dependence of the pH-value form Jdc-was investigated by Thompson and Barlo [35] as
well as Büchler and Schöneich [19]. Based on these results the pH at the steel surface can be
described for the case of a hindered mass transport at the steel surface (a diffusion and
migration controlled transport of OH-) with equation (8) as a function of Jdc. This hindered
mass transport is observed in the case of a coating defect bedded in sand and soil or covered
with calcareous deposits.
dcJppHpH log0 (8)
The R of a coating defect is a key parameter in CP. Based on equation (8) Jdc results in an
increase of the pH (the OH- ion concentration) at the steel surface. These OH
- ions will
migrate in the electrical field and diffuse due to the concentration gradient away from the steel
surface into the surrounding soil. The increase of the pH at the steel surface and the transport
of the OH- ions into the surrounding soil results in a relevant increase of the ion concentration
in soil and, therefore, in a relevant decrease of the spread resistance. An increased Jdc will,
therefore, not only result in an increase of the pH at the steel surface, but also to an increased
migration. This increased migration of OH- ions at increased current densities was identified
6
as the main reason for pH values not higher than 14 at the steel surface even at current
densities as highs as 50 A/m2 [19].
Under the following assumption the spread resistance of a cathode with the shape of a
hemisphere (RH) can be determined with a simple geometrical consideration: The mass
transport moves the generated OH- ions away from the cathode surface. Their concentration
can be calculated in a first approach based on the geometrical dilution with increasing
distance from the cathode surface. Further it is assumed that the conductivity of the dissolved
ions in the electrolyte can be added to the contribution to the conductivity of the OH- ions.
Under these assumptions the change of the soil resistivity can be treated as two parallel
resistors. Hence, the soil resistivity ρx at a distance x from the surface of the hemisphere
according to equation (9) can be calculated based on the dissolved ions ρ in the soil and the
dissolved alkalinity ρpHx. The dilution of the OH- ions due to the transport into increasingly
larger volumes can be calculated from the second term in equation (9).
111 pHxx
11
1222
dxd pH (9)
The resistivity of a soil soaked with NaOH-solution ρpH at the cathode surface at x = 0 can be
calculated according to equation (10) from the pH-value at the cathode surface, the factor ρ0
and the factors a and b.
)exp(0 pHbapH (10)
The integration of the local soil resistance at the cathode surface to the position of the
reference electrode ae results in RH according to equation (11).
pH
pHepHpH
a
pH
a
xH
d
da
dxxd
dxddx
xdR
ee
arctan)/21(arctan
2
22
22
0
2
11
122
0
2
(11)
The integration is performed from 0 to ae with the diameter of the hemispherical cathode d,
the soil resistivity ρ as well as ρpH. For ae usually remote earth with a value of 30 m is used.
From the spread resistance of the hemisphere RH with the diameter d the spread resistance of
a circular defect R with a diameter d is calculated according to equation (12).
2
1114
d
dlRRRR
pHk
HkH
(12)
The procedure for the calculation of R is schematically illustrated in Fig. 2. Based on equation
(8) the pH at the cathode surface is calculated. The geometrical dilution of the OH- ions and
their effect on the spread resistance of a hemispherical cathode is calculated according to
equation (11). Hence, the contribution of the red hemisphere in Fig. 2 is ignored. This
contribution is considered in equation (12) by introducing a correction resistance Rk. It
corresponds to the resistance of an assumed cylinder with the diameter d filled with the pH
that is present at the cathode surface (calculated according to equation (8)). The height of the
cylinder is d and it is corrected by means of the factor lk to fit the empirically determined
7
spread resistances. This procedure numerically suggests a homogenous current distribution
within the coating defect. This is evidently not the case, but it is assumed for the
determination of the threshold values in EN 15280 and ISO 18086, since the measured current
has to be divided by the coating defect surface according to these documents.
Fig. 2: Schematic configuration for the calculation of the spread resistance of a circular
defect with diameter d
For the calculation of the spread resistance the contribution of hydrogen bubbles, of corrosion
products as well as gravity are ignored. Additionally, a possible increase of the electrolyte
volume caused by the metal loss and the contribution of the adjacent coating thickness are
ignored as well. All these assumptions represent the worst possible cases.
Determination of the calculation parameters
The description of the model showed that a series of parameters is required for the calculation
of the critical operation conditions of pipelines. The performed measurements are only
illustrated here for the spread resistance. The complete procedure is given in [6]. A summary
of the optimized parameters is shown in Table 1. Based on their large number and their
partially exponential behavior their optimization was performed in several steps
Table 1: Optimized model parameters as well as the proposed values for further field
optimization. In the case of field optimization different parameters are needed for 3 mm
polyethylene coatings (PE) and 0.5 mm fusion bonded epoxy coatings (FBE).
Parameter Optimized parameters Proposed values for field application
PE FBE
J0 [A/m2] 0.18 0.18 0.18
Kk [V/dec] 0.126 0.126 0.126
pH0 [-] 12.4 12.4 12.4
p [-] 0.5 0.5 0.5
ρ0 [Ωm] 1 1 1
a [-] 12.58 12.58 12.58
b [-] 0.94 0.94 0.94
lk [-] 0.8 0.2 0.2
q [-] 0.4 0.4 0.4
f [mVm2/A] 0.4 0.4 0.4
bu 0 lmax 0
The results of the measurements of the spread resistance RH for hemispheres of 10 mm
diameter are shown in Fig. 3. Additionally the values calculated by means of equations (10)
and (11) are shown. There is a good qualitative agreement between the measured and
calculated values and calculated resistance values are smaller than the measured ones.
8
Fig. 3: Dependence of the spread resistance RH of a hemisphere with diameter 10 mm
from the soil resistivity (33 to 900 Ωm) and Jdc. The open symbols are the calculated
values.
Fig. 4: Dependence of the spread resistance R of a disc shaped electrode with diameter
10mm from the soil resistivity (25 to 750 Ωm) and Jdc. The open symbols are the
calculated values.
9
Fig. 5: Comparison of the current densities measured under laboratory conditions and the
corresponding values calculated by means of Eon, Uac, ρ and the parameters in Table 1.
In Fig. 4 the corresponding results for disc shaped coating defects with a diameter 10 mm are
shown. The calculated values were determined with equation (10) and (11) as well as the
factor lk of 0.8 and equation (12). With the parameters in Table 1 for the relevant current
densities of 1 A/m2 the calculated spread resistance is smaller than the measured ones for all
soil resistivities. Based on these results it can be concluded that the simple model allows for
calculating the spread resistance of coating defects as a function of ρ and Jdc.
The faradic rectification is described by means of equation (4) and the factor f is a key
parameter controlling the processes taking place at the steel surface under CP. In order to
validate the dependencies described by the model and to optimize the factors f and lk,
laboratory investigations were performed with ER-coupons of 1 cm2 at various Eon and Uac in
quartz sand soaked with artificial soil solution according to [19]. In Fig. 5 the results of the
optimization of the parameters listed in Table 1 are shown.
Clearly the model and especially equation (5) are able to correctly describe the behavior of a
steel surface under CP with ac-interference. Considering the coarse simplifications of the
model a very good agreement between the measured and calculated data is found.
In a further step these optimized parameters were applied to the data collected in the field
tests [1, 2]. Despite of a clearly worse correlation than in Fig. 5, there was no case of ac-
corrosion on the coupon where the model predicted non-critical current densities [6]. With
respect to the prediction of the corrosion situation, the model was, hence, always on the safe
side when applying the parameters in Table 1. This again validates the model and the
applicability of equation (5) based on the field data.
Conclusions
Despite various assumptions and important simplifications the model is capable of describing
the relevant influencing factors in ac-corrosion. Based on the optimized parameters in Table 1
the predictions with respect to the occurrence of corrosion were always on the safe side in the
10
case of laboratory as well as field investigations. Based on this it was possible to validate the
model with respect to the geometrical aspects. These are fundamental for the determination of
the corrosion depth at which ac-corrosion is expected to stop.
Validation of the geometrical aspects
Based on equation (5) a decreasing metallic surface A results in very small acceptable Uac. As
a consequence, ac-corrosion cannot be prevented on small coating defects. In contrast, the
increase of A due to the corrosion process should result in a decrease of the corrosion rate
with increasing depth. The current densities will decrease below the thresholds in the
corresponding standard, which then should result in a stopping of the corrosion process. After
the electrical aspects of the model were validated in the previous chapter, these geometrical
effects will be further investigated.
Calculation of the metal surface
The increase of A with increasing corrosion depth is a key effect in the ac-corrosion process.
The description of the dependency of A from the corrosion depth lmax was performed as
follows:
It was assumed that the corrosion site exhibits a ball shape. It can hence be described by
equation (13) based on the quotient q and the diameter of the corrosion site dk. Based on
laboratory and field tests a typical maximum value of q was found to be 0.4 [6].
kd
lq max2
(13)
For the calculation of lmax three cases need to be considered which are illustrated in Fig. 6.
The surface A of the corrosion site exhibits a ball shape from the very early stages on and is
described by equation (14) as illustrated in Fig. 6a. The parameter bu describes the width of
corrosion extending under the coating in the early stages of the corrosion process. In Fig. 6 bu
was chosen as zero and therefore defect diameter d corresponds to the diameter of the
corrosion site dk in the early stages of the corrosion process.
42
22
maxubdlA
(14)
2
max2
max ql
lA (15)
According to the model the ac-corrosion results in a further increase of A according to Fig.
6a. This early stage of the corrosion process is characterized by a q calculated by equation
(13) smaller than 0.4. As soon as q equals 0.4, the situation in Fig. 6b is reached. The further
evolution of the corrosion will result in an extension of the corrosion process under the
coating as shown in Fig. 6c. In this case the A is calculated based on equation (15).
The process described in Fig. 6 has a key implication: Any maximal acceptable corrosion
depth lmax has an associated critical defect diameter dkrit and the associated critical defect
surface Akrit according to equation (16).
11
(a) d
lq max2
(b) d
lq max2
(c) kd
lq max2
Fig. 6: Conditions for the calculation of the surface of the corrosion pit for q=0.4 and
bu=0.a): early stages of the process; b) Limiting condition for the transition from equation
(14) to (15); b) Later stages of the corrosion process.
Akrit corresponds to the surface of the coating defect (not the metallic surface of the corroded
surface A) and is identical to the metallic surface before the corrosion process has started. At
Akrit the acceptable Uac calculated according to equation (5) for the corresponding lmax is
minimal. This geometrical situation corresponds to Fig. 6b, where a maximum lmax is reached
with minimal increase of A.
2
max
2
4u
kritkrit bq
ldA
(16)
Based on the parameters in Table 1 and the acceptable maximum corrosion depth of 2 mm an
Akrit in the range of 1 cm2 is obtained. From equation (16) it follows an Akrit of zero for an
acceptable lmax of zero. Hence the acceptable Uac according to equation (5) will reach zero as
well. This confirms clearly that it is impossible to exclude the occurrence of ac-corrosion.
Interestingly, the critical coating defect size for important ac-corrosion was found to be 1 cm2.
This value as stated in the relevant standards can readily be explained based on equation (16).
There is important consequence of equation (16): The size and the size distribution of coating
defects on pipelines are not known. Hence, the calculation of the admissible Uac must always
be based on Akrit which again is a function of lmax.
Optimization of the parameters in laboratory investigations
These considerations have clearly shown the relevance of q for the assessment of the ac-
corrosion process. It controls Akrit that in turn determines lmax. The key effect in the model is
the decrease of the corrosion rate with increasing A. This aspect was investigated by means of
coated steel plates with an artificially introduced circular coating defect with various sizes
(Fig. 7 left). This type of coupon allows for an increase of A by lateral extension of the
corrosion pit under the coating. In contrast, the rod shaped coupon (Fig. 7 right) prevents a
lateral extension of the corrosion process. In this case an increase of A will not be possible.
Fig. 7: Different geometrical conditions for the evolution of corrosion. Left: Coated plate.
Right: Coated rod.
12
Fig. 8: Corrosion attack on a coupon with 2 mm coating defect diameter. A circle (red) is
plotted into the corrosion profile.
Fig. 9: Evolution of the corrosion process on a rod shaped coupon with 2 mm diameter at
Corrosion attack on a coupon with 2 mm coating defect diameter Eon -1.35 VCSE, Uac-16 V.
The pictures were taken at various intervals. They show the corrosion products and the
metal loss (red line).
13
The exposure of both types of coupons with identical initial metallic surface to identical Eon
and Uac in the same quartz sand soaked with artificial soil solution according to [19] allows
for investigating the evolution of the corrosion process over time. The example shown in Fig.
8 clearly confirms the extension of the corrosion process underneath the coating as shown in
Fig. 7 for the coated plate. This is characteristic for ac-corrosion and was already observed at
the first ac-corrosion leakage in Switzerland [9].
In Fig. 8 the depth profile determined on the plate with 2 mm coating defect diameter is
shown. After 8 months the corrosion depth had only reached 1 mm. Clearly the ball shaped
corrosion pit can be recognized. The results confirm the general applicability of the concept
illustrated in Fig. 6. Hence equation (13) provides a realistic picture of the corrosion shape.
The example for the rod shaped electrode is shown in Fig. 9. The evolution of the corrosion
process could be optically followed by using Plexiglas (PMMA) for embedding of the 2 mm
diameter steel rod. In all experiments corrosion rates significantly larger than 10 mm/years
were obtained [6]. Characteristic for all the tests was the formation of rod shaped corrosion
products that were pushed out of the PMMA.
Conclusions
All the measurements performed clearly show significantly increased corrosion rates on the
rod shaped compared to the plate shaped coupons. The only difference in the experimental
set-up was the lateral limitation of the corrosion in the case of the rods. These experiments
very clearly confirm the expected relevance of A and its increase with progressing corrosion
depth on the corrosion rate. This validates the key conclusion obtained from the model.
The investigation of the corrosion products showed a porosity of 50 to 60%. Within the
corrosion products larger cavities were observed that represent a path for the release of
hydrogen formed at the steel surface. The x-ray diffraction analysis revealed that the
corrosion products consisted of pure magnetite. Furthermore, the resistivity was in the range
of 3 Ωm, justifying the neglecting of the corrosion products in the model calculation.
The characteristic morphology of the corrosion products in Fig. 9 confirms the absence of
soluble corrosion products. The solid state conversion of the passive film during cathodic
polarization as observed by Schmuki et al. [25] can readily explain this behavior. Each
formation of a fresh passive film results in an increase of the volume. The rust formed during
the cathodic dissolution of the passive film through solid state conversion is pushed outward.
The important mechanical forces cause an expulsion of the corrosion products in the case of
the rod shaped coupon and a lift off of the coating in the case of a plate shaped coupon. If the
corrosion process would be a result of soluble corrosion products in the highly alkaline
environment at the steel surface, the formation of a pustule due to precipitation of the
corrosion product in the less alkaline soil would be expected. This confirms that the corrosion
mechanism is not caused by the formation of soluble iron compounds at highly increased pH-
values at the steel surface, as already discussed in [19].
Validation of the model with field data
Within the experiments the dominant influence of the increase of the metallic surface A of the
corrosion site caused by the progressive metal loss was confirmed. The limited testing time of
12 months and the limited precision of the corrosion rate measurement do not allow
demonstrating the complete stopping of the corrosion process once the critical depth was
reached. Based on the available laboratory data it is not possible to exclude that the corrosion
only was slowed down significantly, but continues at a level that still represents a threat to the
integrity of the pipeline. By analyzing the excavation and inline inspection data of the
participating pipeline operators it was possible to apply the model on longer time periods
under realistic interference conditions. The geometry of the corrosion sites observed in the
14
field was compared with the model expectations [6]. Moreover, the expected maximum
corrosion depth based on the model calculations and the parameters in Table 1 was compared
to the values found in the field as shown in Fig. 10. In most cases the depth of corrosion was
overestimated. However, there are some cases with underestimation that need further
discussion.
There is a case of ac-corrosion with a depth of 7.5 mm determined on a coupon, where the
model predicted only 3.1 mm. This coupon was rod shaped (c.f. Fig. 7 right), which did not
allow for lateral growth of the corrosion and provided only limited increase of the surface
with increasing depth. Under these conditions the model expects no decrease of the corrosion
rate with depth, which was indeed confirmed.
There is a series of corrosion depths with 2.5 mm, where the calculated corrosion depth was
clearly smaller. These data were based on internal inspection data and the 2.5 mm are the
resolution limit of the inspection tool. Hence the corrosion depth was indeed smaller or equal
to 2.5 mm.
Fig. 10: Comparison of the corrosion depth measured on coupons and pipelines under
typical operation conditions with the calculated expected maximum corrosion depth.
Further there is a series of data, where no corrosion was expected based on the model
calculation based on the usual operation conditions, but the excavation shows corrosion of up
to 1 mm depth on the pipeline and on coupons. In this specific case the coating defects were
located by means of a DCVG at strongly increased rectifier output in order to obtain a better
resolution. If the Eon and Uac data present during the DCVG are used for the model
calculation, the corrosion depth observed is in line with the model. Considering the very high
corrosion rates in the initial stages of the ac-corrosion process (e.g. Fig. 9) and the time
required for the DCVG on that pipeline the corrosion depth of up to 1 mm/year can readily be
explained.
The calculations demonstrate the applicability of the model for estimating the maximum
corrosion depth. The pipelines were exposed to these interference conditions for longer
periods of time and no leaks were observed in a single case. This validates the model and
15
confirms its applicability for determining the ac-corrosion risk of pipelines under ac-
interference.
Conclusions
The presented model for ac-corrosion is capable of explaining the discrepancy between the
high corrosion rates observed on coupons and the very limited number of damages on
pipelines. The increase of the steel surface due to the corrosion process results with increasing
depth in decreasing current densities. When they reach the thresholds stated in EN 15280 and
ISO 18086 the ac-corrosion process is expected to stop.
The individual parameters were first calibrated in laboratory investigations and then validated
in field tests. The applicability of the model was thus demonstrated. Moreover, the predicted
influence of the surface of the corrosion site on the corrosion rate was confirmed. Hence, all
available information indicates that ac-corrosion will stop at a critical depth. However, based
on all the available data it is not possible to prove this effect due to limited resolution in
corrosion rate measurements.
The presented model is confirming the empirical experience collected in the past 30 years. It
allows, therefore, predicting the critical conditions and optimizing mitigation measures.
Based on the important relevance of the steel surface and hence the corrosion depth, the ac-
corrosion rate must be considered to be of limited relevance. The very high corrosion rates in
the early stages of ac-corrosion decrease rapidly with progressing depth. Hence, it is
impossible to extrapolate a corrosion rate based on an exposure time and a metal loss. All the
available data demonstrate that the assessment of the acceptable interference level must be
based on an acceptable corrosion depth. The discussion of the critical coating defect surface
demonstrates that the assumption of a critical coating defect surface of 1 cm2 already implies
an acceptable corrosion depth in the range of one to two millimeters. The meeting of the
requirements of EN 15280 and ISO 18086 on coupons with 1 cm2 defect surface cannot
exclude higher current densities on smaller coating defects and hence corrosion. As predicted
by the model calculation these small coating defects never lead to perforation of pipelines. It
is expected that they corroded very rapidly in the early stages, but then stop corroding within
1 to 2 mm depth. These considerations clearly show that already the present standards have an
implicitly accepted maximum corrosion depth.
Outlook
The validation of the model allows applying the presented concept for the assessment of the
ac-corrosion risk on pipelines. The numerical description of the relevant influencing factors
offers the possibility to correctly address them and optimize mitigation measures. The present
parameters in Table 1 were optimized based on laboratory investigations and their
applicability to pipelines was demonstrated. However, there is need for further validation and
optimization of the parameters based on currently operated pipelines. The fact that they did
not leak over important periods of time will allow for further optimization of the parameters.
The input of on-potentials, ac-voltages and soil resistivities as well as the threshold values in
EN 15280 and ISO 18086 in the model and the comparison of the calculated maximum
corrosion depth with the wall thickness allows for further validation of the model. Similarly,
the data from inline inspection may be used for this analysis.
Examples of this calculation for an lmax of 5 mm are shown in Fig. 11. Clearly, an important
dependence of the admissible Uac on ρ and Eon is found, demonstrating that it is impossible to
define generally valid interference levels. The data in Fig. 11 allow for determining critical
sections of the pipeline system as well as the development of mitigation strategies. It has to be
pointed out, that depending of the various factors the highest corrosion risk is not necessarily
linked to the highest Uac.
16
(a)
(b)
Fig. 11: Admissible average Uac as a function of the average Eon for various soil
resistivities ρ and an acceptable lmax of 5 mm calculated with the parameters in Table 1
for: a) PE and b) FBE coatings.
With this approach shown in Fig. 11 it was already possible to demonstrate relevant
differences between fusion bonded epoxy coatings (FBE) with thickness in the range of 0.5
mm and three layer polyethylene coatings (PE) with thickness in the range of 3 mm. In the
case of FBE the corrosion products fracture the coating and the extension of the corrosion
process underneath the coating is very limited. Therefore, the defect diameter d is increasing
with the diameter of the corrosion site dk. The corrosions process is, therefore, controlled by
the situation shown in Fig. 6b in the case of FBE. This has relevant implications for the
corrosion process according to the model. Small coating defects result in high current
densities and high corrosion rates. Instead of a fast stopping of the corrosion (as expected for
PE), the coating defect diameter grows with increasing corrosion depth and allows for further
extension of the corrosion process. This is expected to result in faster perforation of the
pipeline in the case of FBE compared to PE coating, since the corrosion process is always run
at the critical defect diameter.
In contrast, the PE coating is mechanically more robust. The experience shows that it will be
lifted off the steel surface through the mechanical pressure of the corrosion products, while
maintaining the original coating defect diameter. Based on these considerations different
parameters are required for addressing the ac-corrosion risk on PE and FBE coated pipelines.
This is addressed with the two different parameter sets proposed in Table 1.
Acknowledgement
This work was only possible thanks to the support by the DVGW, ENBW Regional AG,
Open Grid Europe GmbH, MERO Pipeline GmbH, MVV Energie AG, Westnetz GmbH,
ONTRAS - VNG Gastransport GmbH, GASCADE Gastransport GmbH and Thyssengas
GmbH. A special thank goes to Prof. Dr. H.P. Büchler for the algebraic solution of the
integral in equation (11).
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19
Annex: Abbreviations
A Surface of a corrosion site
Akrit Critical coating defect surface
ae Distance of the reference electrode from the surface of a hemispherical electrode.
Usually remote earth with a value of 30 m is chosen.
a Parameter for equation (10) according to Table 1
b Parameter for equation (10) according to Table 1
bu Width of corrosion extending under the coating in the early stages of the corrosion
process at q>2*lmax/dk
d Diameter of the coating defect
DCVG Direct Current Voltage Gradient measurement
dk Diameter of the corrosion site
dkrit Critical diameter of the coating defect that results at a given lmax in the smallest
Uac
DVGW German Technical and Scientific Association for Gas and Water
E0 Equilibrium potential of the hydrogen evolution
EIR-free Electrode/soil potential without IR-drops caused by electrical currents. It
corresponds to the arithmetic average of the electrode/soil potential recorded over
at least 0.1 seconds
ΔEF Contribution of the Faradic rectification to EIR-free
Eon On-potential measured with a distance ae of the reference electrode to the surface
of a hemispherical electrode
EH Electrode/soil potential of hydrogen evolution without IR-drops caused by
electrical currents.
f Factor for calculating the ΔEF
FBE Fusion bonded epoxy
Jdc Protection current density
J0 Exchange current density of the hydrogen evolution
Kk Tafel slope of the hydrogen evolution
lk Correction factor for equation (12) according Table 1
lmax Maximum corrosion depth
p Parameter for equation (8) according to Table 1
PE Polyethylene
pH0 Parameter for equation (8) according to Table 1
PMMA Plexiglas
q Quotient describing the geometry of the corrosion site
R Spread resistance of a circular coating defect with diameter d
RH Spread resistance of a hemispherical electrode with diameter d
Rk Correction resistance for equation (12)
Uac ac voltage measured with a distance ae of the reference electrode to the surface of
a hemispherical electrode
x Distance from a hemispherical electrode
ρ Soil resistivity
ρx Soil resistivity at distance x from a hemispherical electrode
ρpH Resistivity of clean soil soaked with NaOH of a given pH
ρpHx Resistivity of clean soil soaked with NaOH of a given pH at distance x from a
hemispherical electrode