Abstract :The aim of this fieldwork was the study of the effect of 50Hz AC, induced by high-voltage power lines, on the cathodic protection system of a natural gas pipeline. The effectiveness of cathodic protection was checked through in situ long-term monitoring and analysis of pipeline electrical parameters. The results gave an insight into the problems of the cathodic protection system operation, caused by AC interference. An AC and DC potential interdependence was observed, that previously has hardly been reported, and was scrutinized in relation to cathodically protected pipelines. The effects of the AC-interference and low frequency DC potential fluctuations, as well as the potential deviations from the protection potential, are examined. These phenomena are associated with corrosion susceptibility and difficulties in obtaining reliable cathodic protection measurements.Introduction :

Pipelines sharing close rights-of-way with high-voltage power lines are subjected to the influence of the electromagnetic field that is created by operational and fault currents from the high-voltage lines. Installing pipelines in energy utility corridors containing high-voltage AC transmission lines exposes the pipelines to induced AC voltage. This effect can be caused by an imbalance between phases in the transmission system and/or by high voltages near transmission tower grounding systems, which result from lightning strikes and/or phase faults. Thus, corrosive and damaging AC voltages may be induced in the pipelines (Smart et al., 1999).

Perforation of high-pressure gas pipelines, due to the influence of alternating current corrosion (AC-corrosion), has been reported, even when the lines were “properly” cathodically protected. Even in cathodically protected structures, the corrosion rate may exceed 1mm/yr (Heim and Peez, 1988; Meier, 1989; Bindschedler and Stalder, 1991).

In the past, the AC-corrosion effect on buried metallic pipelines was considered negligible, because the use of old-type coatings (e.g. coal bitumen, etc.) resulted in a rather low insulation resistance towards ground. In fact, this type of insulation, which was made of porous materials, mitigates the induced voltage over the metallic buried pipeline in the presence of crossings or extended parallelisms with high-tension electric lines. The presence of the distributed porosity causes a uniform AC dispersion to ground, which causes a very low AC voltage between pipe and remote earth. The AC current density for such a uniform porosity of the coating along the pipeline is practically negligible. However, when the latest generation of coatings (e.g. three-layer polyethylene) are used, the electric insulation values of the pipeline against ground can become very high (e.g. from 5,000-50,000 Ohm × m2 for bitumen coating to 300,000-1,000,000 ohm × m2 – and even more – for polyethylene). Consequently, in well-coated lines, the risk is greater because of the higher AC current density (Heim and Peez, 1988; Devay et al., 1964). In polyethylene-coated lines, AC can be discharged to ground through tiny coating defects. In this case, the available surface is very small, and this results in a higher AC current density under the same environmental conditions. From the corrosion standpoint, AC current densities over 20A/m2 are generally considered to be hazardous (Heim and Peez, 1992; Peez, 1993; Beuth, 1992). It has also been reported that the areas most corroded by induced AC on pipelines are those with a surface of approximately 1cm2 (Heim and Peez, 1992).

The AC-interference produces cathodic current and IR-free potential fluctuations that have the same frequency as the AC excitation waveform (i.e. 50Hz). Recently reported investigations suggest that AC-corrosion is better correlated with the corresponding IR-free potential excursions from the protection potential (Pourbaix et al., 2000; Schöneich and Meliš, 2001). Laboratory investigations have revealed shifts of the DC potential and an increase of the cathodic current density, under AC-interference (Devay et al., 1964; Chin and Venkatesh, 1979; Chin and Sachdev, 1983). Similar conclusions also were reported by investigating AC-interference in a pipeline case (Kajiyama and Nakamura, 1999). Low frequency (1cycle/min) DC fluctuations, caused by 50Hz AC effects, were detected near the active/passive transition potential of carbon steel in sulphate solution, and were attributed to periodic formation and breakdown of passive film on the steel surface. Nevertheless, the mechanism of these oscillations was not clarified (Chin and Venkatesh, 1979). Generally, the connection between varying AC-induced voltage and low frequency DC potential fluctuations in cathodically protected pipelines is not well understood. Besides, low-frequency potential deviations can be accompanied by increased corrosion rates if the fluctuating potential occasionally deviates from the protection potential. In such cases, protection is insufficient, if only the average value of the potential over time remains more negative than the protection potential (Heim, 1982; CENELEC, 2000).

Distributed-source transmission line (DSTL) theory predicts that the potential variations are the largest at the ends of the pipeline. At opposite ends of the pipeline the potential fluctuations are out-of-phase in the case of constant electromagnetic field (Boteler and Seager, 1998; Taflove and Dabkowski, 1979). In addition, the presence of alternating voltage and current on buried metallic pipelines may cause malfunctioning and even breakdown of the unsuitably equipped transformer/rectifier (T/R) units of an impressed current cathodic protection. The most frequent malfunctions of such cathodic protection T/R units, which are caused by the presence of AC on the pipeline, are:

circulation of unidirectional current in the system pipeline-anode bed; limitation of the current supplied by the rectifiers; modification of the reference potential set on the automatic rectifiers (Gentile et al.,

1997); rectification of AC voltage through T/R unit in cases where there is no special AC-filter

diverting AC around the rectifier diodes (Ames, 1993).

Such problems also make it difficult to specify an optimum protection potential for the pipeline. This is especially important when the protection potential must also not exceed over-protection potential limit to avoid possible risk of phenomena such as stress corrosion cracking, hydrogen embrittlement and cathodic disbonding of the coating.

Experimental

The surveyed pipeline was a part of a natural gas transmission main installed between 1993 and 1995 and buried at a depth of 1 to 2m in the ground. The pipeline (API 5L X-60) was 762mm (30in.) in outside diameter, with a wall thickness ranging from 9.52mm to 15.6mm.

The line was coated with a factory-applied three-layer extruded polyethylene coating, and heat-shrinkable polyethylene-compatible sleeves covered every welded seam (applied in the field). Valves were coated with a coal tar polyurethane coating (Isverflex).

Impressed current cathodic protection was applied by means of one T/R unit in potential control mode. The ground-bed comprised magnetite (Fe3O4) anodes. The T/R unit was equipped with a special low-pass AC-filter, to prevent AC voltage on the pipeline from interfering with T/R unit operation. The pipeline under examination was 45km in length, and was electrically separated from the rest line by buried isolating joints at both ends (at chainage 2.3km and 47.3km). The pipeline was not affected by DC stray or telluric currents and there were no electric trains in the area. Long-term AC-interference on this pipeline was caused by induction from the electromagnetic fields generated by electric power transmission lines (eight 150kV and two 380kV lines). A map displaying the high-voltage lines and the pipeline routes is presented in Figure 1.

The pipeline was equipped with 44 cathodic protection (CP) measuring posts, which allowed measurements of potential and current. This included 18 test posts, named “KG”, which allowed measurements both on pipe and on grounding wires that had been installed at different sites along the pipeline. The problem of the proximity of high-voltage lines had been taken into account during the design phase of the pipeline construction. After installation of the pipeline, grounding wires made from hot-dipped galvanized steel had been installed at selected locations in parallel with the pipeline at a minimum distance of 0.2m from it (Figure 2).

The external diameter of the earthing wires was minimum 12mm with 70μm zinc layer thickness, corresponding to 500g/m2. Table I shows typical grounding installation data, i.e. the chainage, the respective earthing wire length, the measured AC resistance (with an AC resistance meter), the calculated DC resistance of earth wire to remote earth and the corresponding average soil resistivity at every earthing site.

The number and location of the grounding installations were selected by taking into account a previous proximity effects study, where induced levels of long- and short-term overvoltage were calculated for the entire length of the pipeline.

Semiconducting over-voltage arresters were connected between pipeline and earth (grounding wire) in the 18 “KG”, type test posts. In principle, these arresters consisted of a set of anti-parallel thyristors that were activated

by an AC potential between the pipe and earth. When the arrester is active, the alternating current is dispersed from the pipeline through the arrester into the ground, so that the AC voltage of the pipeline is reduced. The actual activation voltage of the arresters was measured in the range from 22 to 26V. During the low load time, i.e. below activation voltage, the arrester is inactive, behaving like an open switch. Technical data for the arresters are given in Table II.

As shown in Figure 3, when the arrester is activated, a voltage drop in the forward direction, Uf, is present across the arrester. The zero flow of the discharge current deactivates the over-voltage arrester. The arrester is reactivated in the next half cycle of the over-voltage at the voltage Un. The over-voltage arrester is not activated, if the amplitude of the over-voltage is less than Un. However, on cathodically protected pipelines, part of the protection current can be absorbed by activated arresters. Arrester installations may have an option of adding into the current path from the earthing installation to the pipe from one to three additional diodes, thus compensating for the DC consumption. Notwithstanding, by the aforementioned method, DC compensation can only be stepwise, and is thus not fully controlled, since it is limited by the 0.7V threshold voltage of each thyristor diode (Paul and Prinz, 1997; Paul and Schöneich, 1997).

The purpose of the over-voltage arresters, for protecting operation personnel from electric shock and for the protection of the pipeline from surges in protection current (short-term over-voltages) will not be examined in this paper. With regard to cathodic protection, field experience indicated that the overvoltage arresters did not satisfy the desired functionality for the cathodic protection system, as it will be shown below.

As for the field measurements, the recording of AC and DC voltage was performed using “Weilekes Elektronik Minilog 512Kbyte” on the 30V AC range (input resistance: 1MΩ) and 3V DC range (input resistance: 2MΩ, attenuation at 50Hz: 60dB) respectively. This recorder was a two-channel data logger, with a built-in AC-filter, which allowed simultaneous recording of DC and AC voltage. The AC-filter of the data logger prevented the DC voltage recordings from being falsified by AC interference. The sampling rate was set at 2 sec. A “Fluke 87-III true rms multimeter” was also used when making the reported measurements.

“On/off” switching was achieved by a “Weilekes Syntakt” time relay. Care was taken to connect the time relay in such a way that, during the “off” period, the pipe was completely isolated from any equipment of the T/R unit. Switching intervals were set at 12 or 27sec “on” and 3sec “off”.

The effectiveness of the cathodic protection system was estimated by measurements of the “on” potential, the “off” potential, and the AC proximity voltage.

The data loggers were utilised at selective test posts. Both AC and DC voltage across pipe and grounding wire were recorded at test posts “KG 47.3km” and “KG 14.3km”. DC and AC potentials versus Cu/CuSO4 reference electrode were recorded at test posts “KG 2.3km”, “KG 22.6km” and at the T/R unit. Additionally, the output direct current supplied by the T/R unit was monitored by means of a data logger connected to the terminals of the T/R unit ammeter. The sampling rate of all data loggers was set at 2sec, so that the potential value during “off” periods was recorded.

The line current measurements were carried out using a microvoltmeter with a 50Hz attenuation of 90dB. By means of this instrument, voltage drops were measured at pipe sections of 50m in length. Given that the resistance value of every 50m-pipe section was known, the line current was easy to calculate.

The T/R unit station was located at chainage 40.1km. It was in the area where the highest AC proximity voltage was usually measured. Typical operating parameters for the T/R unit are shown in Table III.

Results and discussion

The time dependence of AC and DC voltage across pipe and earth wire, monitored at measuring post “KG 47.3km”, is presented in Figure 4.

The DC voltage across pipe and earthing wire is defined as follows:(see equation 1)

where:

ΔE: DC voltage across pipe and grounding wire;

Ep(on): pipe-to-soil “on” potential versus Cu/CuSO4 reference electrode;

Eg: grounding wire potential versus Cu/CuSO4 reference electrode

The test post “KG 47.3km” was situated at one end of the pipeline section, near the isolating joint, where the highest long-term AC induced voltage was regularly recorded. An over-voltage arrester was permanently installed there. Various cathodic protection measurements along the pipeline, as well as simultaneous data logger monitoring both at the ends and at intermediate locations of the pipe, confirmed that the arrester at test post “KG 47.3km” was active only periodically during the measurement period. As shown in Figure 4, time intervals of active and inactive arrester were interchanged alternately. The arrester was active during 33 per cent of the entire data logger monitoring time, i.e. for 31.6 hours in 95.5 hours of total recording time. Additionally, during inactive arrester periods, the pipeline exhibited an AC voltage ranging from 12 to 25V, whereas during active arrester periods the AC voltage was less than 10V.

It was apparent from the recordings that the over-voltage arrester allowed pipeline to attain a high AC-induced voltage level (up to 25V) for long time intervals. Nevertheless, corrosion caused by alternating currents (AC-corrosion) on cathodically protected pipelines has been reported, even at voltage levels below 5V (Heim and Peez, 1992). It was, therefore, imperative to reduce the AC voltage on the pipeline.

From Figure 4, it is also evident that DC voltage-time dependence correlated with the corresponding AC voltage-time dependence. It was noted that the activation periods of the over-voltage arrester, which occurred at approximately 25V, were delimited by abrupt AC potential changes. During the periods when the arrester was active, a considerable shift of the DC voltage between pipe and earth towards 0mV was recorded, suggesting that direct current was flowing from the grounding wire to the pipe, passing through the arrester. Thus, the potential of the pipe and potential of the earth wire tended to achieve similar values.

Simultaneously, as is shown in Figure 5, the total direct current fed by the T/R unit was higher during active arrester periods.

The results suggest that direct current was being conducted by the active arrester and was consumed from the grounding wire. Besides, the direct current leak during activation time of each arrester was of the order of a few mA. In the case of very well insulated pipelines, direct current of that order of magnitude is a large portion of the total protection current supplied from T/R unit, and thus it will affect the cathodic protection function. Hence, cathodic protection of the steel pipeline was affected adversely by arrester activation. By contrast, the same current leak would be negligible in the case of porous coatings, because the cathodic protection current usually would be much larger. As there is a general tendency to construct well-insulated pipelines, the necessity for manufacturing devices with a very high DC blocking capacity is becoming increasingly evident.

Additionally, it is important to point out that during periods when the arrester is inactive, or under conditions of high AC voltage (especially between 20 and 25V), a similar trend of the DC voltage existed, though to a lesser degree. In particular, the DC voltage exhibited a tendency to move towards less negative values, as is shown in Figure 4. Simultaneously, the direct current supplied from T/R unit increased with AC potential, as can be seen in Figure 5. These results imply that AC-interference has affected the DC potential values. The mechanism of this DC potential change may be attributable to the depolarisation effect of AC on the cathodic polarisation characteristics of steel, which has been observed in experiment tests (Devay et al., 1964; Chin and Venkatesh, 1979; Chin and Sachdev, 1983; Gummow et al., 1999).

The corresponding time dependency of the pipe-to-soil potential at T/R unit site, with conditions at test post “KG 2.3km”, is shown in Figure 6.

A potential instability is evident, although the T/R unit did not only operate potentiostatically; it was also shielded from AC-interference by means of a special AC-filter. A comparison between plots shown in Figure 4 and Figure 6 reveals that the potential measured at the T/R unit site followed the trend of the DC voltage that was measured at test post “KG 47.3km”.

The pipe-to-soil potential variations at the T/R unit location (close to one end of the pipeline), and at test post “KG 2.3km” (at the other end of the pipeline), had reverse correlation patterns (i.e. they were out-of-phase).

Additionally, the potential variations recorded at “KG 2.3km” were intensified approximately five times, compared with the simultaneous potential variations at T/R unit location. This result was in agreement with DSTL theory, according to which the potential fluctuations at opposite ends of the pipeline are out-of-phase in the case of constant electromagnetic field excitation. However, this result may be considered to be coincidental in the present instance, as the theory predicts that, for example, when the electromagnetic field has a linear relationship with pipeline chainage, the potential fluctuations at opposite ends of the pipeline would be in-phase. The same theory also predicts that the largest potential variations would occur at the ends of the pipeline (Boteler and Seager, 1998; Taflove and Dabkowski, 1979).

Laboratory investigations have shown that the magnitude of the AC voltage level affects the DC potential (Devay et al., 1964; Chin and Venkatesh, 1979; Chin and Sachdev, 1983). It is, therefore, expected that, when a varying AC voltage is present, it could affect analogously the average DC potential. Thus, slow deviations in the average pipe-to-soil potential, such as those depicted in Figure 6, could be attributable to the respective AC voltage variations. Those pipe-to-soil potential deviations may become significant near the ends of the pipeline, as has been described previously.

Concerning the total cathodic protection current supplied from T/R unit, the average value was approximately 25mA. During periods when the arrester was active, the current values increased to more than 30mA, whereas, during intervals when the arresters were inactive, an average value of 20mA was found. However, during the inactive periods, occasionally ascending AC-induced potentials were accompanied by a corresponding increase in the output of direct current supply from T/R unit, as is shown in Figure 5.

Furthermore, during active periods, the IR drop (i.e. the difference between the “on” and “off” potentials) had a tendency to decrease. “Off” potential values recorded at the T/R station were more negative (approximately 60 to 100mV in average) compared with those during inactive arrester periods. Representative results showing this effect are displayed in Figure 7.

This shift to the negative direction may be attributable to a falsification of the measured “off” potential, due to cell current flow from the earthing installation to the pipeline. The cell currents were caused by the active arrester, which was DC conductive. Additionally, the following effect intensified the consequences of the aforementioned cell currents. The DC potential of the earth wire, which was made from galvanized steel, was relatively electropositive, after a short time in service. For instance, the earthing wire potential at test post “KG 47.3km” was 650mV versus Cu/CuSO4 reference electrode, whereas normally it should be expected to be between 900mV and 1,100mV, i.e. in the range of the natural potential of zinc. This result may be attributable to corrosion of the grounding wire zinc plating and/or by the presence of iron-zinc alloy on the wire surface, due to inefficient galvanizing because of the so-called Sandelin effect (Pistofidis, 2001).

A closer examination of the data in Figure 7 revealed that the short-term voltage peaks tended to occur during periods when the arrester was active. In addition, scrutiny of the data in Figures 4-6 indicates that the DC voltage fluctuations had a tendency to increase with AC voltage.

As is shown in Figure 8, DC/AC potential fluctuations of low frequency also were recorded, especially at high AC voltages.

It seems unlikely that these fluctuations were caused by inefficient AC-filtering in the data logger. Hence, these data were assumed to be real, though the DC potential fluctuations had significant amplitude only at “KG 47.3km”, while at more distant locations they exhibited only a very low amplitude. A possible explanation may be that the over-voltage arrester may have been in a semi-active condition and that resulted in the observed interference effect. As the data logger sampling rate was relatively high (2sec), an aliasing effect is possible, as a result of which the frequency of the oscillations could not be determined.

Line current values for the entire AC-subjected pipeline length are plotted in Figure 9. As is evident, no clear conclusions could be drawn from the results, though the corresponding AC proximity voltage was relatively low during line current measurement, as is shown in Figure 10.

An additional consequence of AC-interference is the inability to acquire meaningful line currents. Hence, it is almost impossible to arrive at any conclusion concerning the effectiveness of a coating system on a pipeline section without increasing considerably the current output of the cathodic protection T/R unit.

It is also worth mentioning that the natural potential values of the depolarised AC-affected pipeline were more negative than normal. A similar phenomenon was observed in laboratory investigations of the AC effect on the corrosion of steel (Devay et al., 1964), and the effect could only be attributed to the non-symmetric influence of AC interference on the cathodic and anodic electrochemical reactions (Chin and Fu, 1979).

The above analysis shows that there is a definite need for the installation of devices that have been designed specifically to reduce AC voltage on a pipeline without affecting the cathodic protection function.

Conclusions

Long-term monitoring and analysis of the electrical characteristics of a cathodically protected natural gas pipeline which had been equipped with semiconducting over-voltage arresters and was subjected to a 50Hz AC-interference, led to the following conclusions:

1. (1) During inactive over-voltage arrester periods:

the pipe-to-soil potential, the DC voltage level across pipe-to-earth, the protection current and the DC voltage fluctuations indicated that a significant contribution from induced AC voltage could occur;

slow deviations of pipe-to-soil potential were recorded, which influenced adversely the efficiency of the cathodic protection system;

low frequency DC/AC voltage oscillations were recorded close to the over-voltage arrester activation voltage, and this behaviour was attributed to a semi-active stage in the arrester function;

a secondary effect of the AC interference was the difficulty of acquiring accurate cathodic protection measurements (e.g. line current and potential);

high AC-induced voltages were present, implying increased AC-corrosion risk.

1. (2) During active over-voltage arrester periods:

the arresters conducted DC, thereby interfering with the operation of the cathodic protection system and reducing its effectiveness;

although the AC voltage was lowered by the arrester activation, it was still sufficiently high for AC-induced corrosion attack to take place;

the difficulty in acquiring reliable cathodic protection measurements was intensified.

Table I Pipe-line grounding system data

Table II Technical data of semi-conducting over-voltage arrester

Table III T/R unit typical operating parameters

Figure 1 Map displaying the high-voltage lines and the pipeline region of AC-interference by electromagnetic field

Figure 2 (a) diagram of over-voltage arrester and earth wire connections at the CP test post and (b) earthing wire installation detail (cross-section)

Figure 3 50Hz AC voltage waveform for activated overvoltage arrester

Figure 4 Time dependence of DC and AC voltage across pipe-earth wire at test post “KG 47.3km”

Figure 5 Time dependence of total direct current supplied from T/R unit

Figure 6 Time dependence of “on” DC pipe potential vs Cu/CuSO4 reference electrode at T/R unit station site (40.1km), inset shows the respective time dependence of “on” DC pipe

potential at test post “KG 2.3km”

Figure 7 Time dependence of “on/off” pipe-to-soil vs Cu/CuSO4 reference electrode at T/R unit station site (40.1km)

Figure 8 Fluctuations of DC and AC voltage between pipe and earth wire at test post “KG 47.3km”

Figure 9 Line current versus chainage

Figure 10 AC potential versus chainage