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, markus.buechler@sgk.ch

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

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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].

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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

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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

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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.

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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.

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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

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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).

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(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.

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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).

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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

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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