EVALUATION OF ALLOY 625 STEEL BIMETALLIC PIPE

FOR PETROLEUM SERVICE

Dr. S. Mark Wilhelm and Dr. Russell D. Kane Cortest Laboratories, Inc.

11115 Mills Road - Suite 102 Cypress, Texas 77429

Dr. S. Matsui, Dr. T. Yoshida, Mr. T. Iwase Kawasaki Heavy Industries, Ltd. 1-l Higashikawasaki-cho 3-Chome

Chuo-ku, Kobe 650-91 Japan

ABSTRACT

The results of laboratory investigations are presented which were conducted to better define the serviceability of bimetallic pipe manufactured by the thermohydraulic (tight-fit) method for Mobile Bay service. Both Alloy 625/API X-65 and Alloy 825/API X-65 tight- fit pipe (TFP) were evaluated under conditions of (1) standard corrosion test environments to evaluate the metallurgical conditions of the corrosion resistant alloy (CPA) liner tubes, (2) long term, full scale TFP exposure under a simulated Mobile Bay production environment containing high levels of hydrogen sulfide and carbon dioxide, and (3) hydrogen permeation experiments designed to examine potential effects of CPA liner collapse from hydrogen produced by corrosion on the ID and/or cathodic protection on the OD. The results of the study indicate that bimetallic TFP exhibits acceptable metallurgical condition of the CPA liner materials. Under the simulated Mobile Bay production environment, TFP exhibited good resistance to general corrosion, SCC and liner Collapse. Hydrogen permeation tests indicate that very conservative estimates of service life liner collapse from interfacial hydrogen pressure are in the range of 150 years to over 800 years depending on the use condition. For all practical purposes liner collapse from hydrogen is not a limiting factor for TFP flowline appliCatiOnS.

Introduction

The application of corrosion resistant alloys (CPA's) for service in offshore oilfield production environments has been a topic of significant interest. It involves both performance and economic questions. For example, one must consider the service performance (corrosion, pitting and stress corrosion cracking) of the materials. In addition, their cost effectiveness relative to the more conventional methods of corrosion mitigation such as the use of inhibitors and coatings must be evaluated.

Two common approaches for applying CPA materials are (1) the use of solid (monolithic) CPA components and (2) the use of bimetallic

Superalloys 718,625 and Various Derivatives Edited by Edward A. Lmia

The Minerals, Metals & Materials Society, 1991

771

or clad CPA components. In many applications, economic or engineering limitations dictate that solid CPA approaches are not always feasible. Presently accepted examples of this are clad wellhead components and process vessels which utilize CPA's on the I.D. that are installed by either weld overlay welding or HIP (Hot Isostatic Pressing) cladding.

The purpose of the present study is to evaluate the serviceability of bimetallic tubular components manufactured by a thermohydraulic (tight-fit-pipe - TFP; tight-fit-tube - TFT) manufacturing process for severe sour service in Mobile Bay. This method involves the assemblage of a CPA liner tube inside of a finished steel outer tube or pipe. The finished tubular thereby has the load carrying capability of the steel pipe along with the internal corrosion resistance of a CPA pipe but has a lower overall cost than solid CRA.

In the present study, an experimental program was conducted to evaluate two key aspects of TFP performance as they relate to application of this material system -for severe sour service in Mobile Bay offshore flowlines. They are (1) the resistance of TFP to the sour production environment in terms of corrosion and stress corrosion cracking (SCC), and (2) the resistance of TFP to collapse of the CPA liner as a result of hydrogen build-up at the interface. This hydrogen is the result of corrosion on the ID and cathodic protection of the steel outer pipe in seawater.

Backsround

Thermohvdraulic (Tiqht-Fit) Fabrication Method

As shown in Figure 1, the thermohydraulic method of fabrication used in TFP involves the hydraulic expansion of a CPA liner tube into a heated steel outer pipe. As a result of this expansion and subsequent cooling to room temperature, a compressive residual stress is induced into the liner tube. This compressive residual stress puts the CPA liner and steel outer pipe into intimate contact creating a tight fit or mechanical bond. This process is contrasted with a metallurgical bonding process for bimetallic construction whereby the two materials must be fused by welding, extrusion or diffusion bonded by HIP methods.

In the case of metallurgical bonding of CPA materials and steels, the thermal treatments necessary for the steel outer pipe can result in a non-optimum condition of the stainless or nickel alloy liner tube. Such conditions can promote localized corrosion and SCC of the CPA material. Similarly, the thermal heat treatments commonly used for stainless and nickel alloys can result in inferior mechanical properties in the steel outer pipe. However, for TFP, the thermomechanical bonding process delivers both the CPA liner tube and steel outer pipe in the final optimized condition which should yield superior performance of the overall product.

To complete the TFP fabrication process, the CPA liner and outer pipe must be seal welded as shown in Figure 2. Furthermore, for flowline applications, TFP can be joined by conventional field welding techniques as shown in Figure 3. Therefore, one of the more critical aspects of TFP performance is the corrosion behavior in this welded region in the severe conditions commonly found in

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Mobile Bay produced fluids which may contain high levels of hydrogen sulfide, carbon dioxide and brine.

Potential Liner Collapse Problm

One of the continuing debates in the oilfield has been the potential collapse of CRA liners in bimetallic tubulars. experience,

Early with downhole bimetallic tubulars revealed that liner

collapse could be a potential failure mode in sour gas wells. This experience was for the most part obtainable using loose liner bimetallic tubes. In these cases, atomic hydrogen produced on the backside of the tubing by sulfide corrosion may diffuse into the material and accumulate at the CBA/steel interface. Continued accumulation has been shown to result in liner blistering and collapse of some bimetallic tubular systems in laboratory tests even when metallurgical bonding techniques have been utilized. [l]

However, in the previously cited study, one of the materials not found to exhibit liner collapse from hydrogen accumulation even under severe conditions of hydrogen charging was pipe manufactured by the thermohydraulic method. Even though these tests demonstrated the superiority of TFP over other available methods, continued discussion has taken place regarding its behavior from the standpoint of liner collapse.

long term

To understand the phenomenon of liner collapse from hydrogen accumulation more fully, it is necessary to examine the stresses that act on the liner tube. Liner collapse should occur when the collapse stress (SC) is equal to the stress due to hydrogen pressure at the CPA/steel interface (S,,) plus the fit in stress in the liner resulting from thermohydraulic processing (Sf) which may be written

s, = s, - s, (1) Sh = s, - s,

The problem is then defined as determining the hydrogen pressure (P,) as given by the following equation:

Where D, = I.D. of outer pipe and t, = the thickness of the CPA liner tube. Combining the equations then yields

p h

= (Sf - S,) (2tJ

01

Substituting appropriate values for typical TFP products as given below: S, = 345 MPa for Alloy 625

S, = 242 MPa for Alloy 825 DL = 230 mm t, = 3.0 mm S, = 95.2 MPa

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results in hydrogen pressures of 11.5 MPa and 8.8 MPa for Alloy 625 and Alloy 825, respectively.

The next requirement is to know the flux of hydrogen atoms entering the interfacial volume between the outer pipe and liner tube. This can be obtained through use of Fick's first law of diffusion [2]

Flux (H,) = $ = (2) (F) (D) CC,)

X (4)

where, z = constant, F = Faraday's constant (96,485 C/mole electron), D = Diffusion constant, concentration,

C, = subsurface atomic hydrogen X = diffusion distance, and im = steady state

hydrogen permeation current. The value of ioo can be obtained through experimental measurements of hydrogen charging efficiency at various charging currents (i,,,) and is given by the following equation:

190 a=- 1 ch

(5)

The efficiency (QI) is a function of environment (water chemistry, cathodic protection, temperature, etc ) and is empirically determined.

By application of Fick's law relative to varying diffusion distances in combination with the preceding equation, the results of laboratory hydrogen charging experiments can be related to actual TFP performance as follows:

where ioo = current related to the hydrogen flux into the material (liner Opr outer pipe), X = membrane thickness in laboratory experiment, X = outer pipernor liner tube thickness.

For conditions of cathodic protection on the O.D. of the TFP outer pipe in slowly moving seawater, PA/cm'.

ich values are commonly 5 - 6 [3] In the case of corrosion on the CPA liner tube, it is

assumed that a tolerable general corrosion rate of 2 mpy could be sustained. This relates to an ic, value of 5 pA/cm2.

The im values can be converted to hydrogen mole fluxes (n,) as shown below:

nf = L (96487)(2)

(7)

Assuming the ideal gas law for hydrogen, pressure increase (P,) can be calculated

the rate of hydrogen gas

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p = &(R) (T) (8) I V

where R = gas constant (0.082 L atm/K mole), T = gas temperature (K) and V = volume under 1 cm2 of pipe wall.

Technical Approach

In the present study, a series of tests were conducted to evaluate the overall performance of TFP for Mobile Bay flowline applications. These tests included (1) laboratory coupon studies in standard environments used to evaluate the metallurgical condition of CRA materials, (2) full scale tests of two TFP CRA/ steel bimetallic pipe systems in a simulated Mobile Bay production environment and (3) hydrogen permeation tests under conditions of simulating hydrogen charging that could take place on the flowline I.D. from CRA corrosion in produced fluids and on the flowline 0-D. from cathodic protection in seawater.

Experimental Procedure

Standardized Tests of Allov 625 Liner Materials From Fabrication Weldments

TFP weldments were fabricated in the manner shown in Figure 3 using an Alloy 625 liner tube and API X-65 steel outer pipe. The weldment consisted of a single triple pass seal weld sample. A sample ring of the liner material was provided by machining off the steel outer pipe after welding. It consisted of a ring of welded Alloy 625, 127 mm in length by nominal 103 mm I.D. and 2 mm wall thickness. The coupons were then machined from the ring sample as shown in Figure 4 and designated with the following code: LS = liner sample (parent metal), BW = butt weld and OL = overlay (seam weld).

The standard test methods used to evaluate the materials are given in Table 1. They include: (1) ASTM G-48 (ferric chloride test), (2) ASTM G-28 (ferric sulfate/sulfuric acid test), (3) ASTM A-262C (nitric acid test) and (4) HCl test. As indicated in Table 1, these tests were utilized to evaluate potential pitting, intergranular corrosion and fusion line corrosion in the Alloy 625 liner tubes.

Full Scale TFP Pipe Sample Tests

Two full scale TFP pipe sections were evaluated in a simulated Mobile Bay flowline production environment. The test pieces were fabricated from TFP into pressure vessels by installing Alloy 625 plugs in both ends of the pipe as shown in Figures 5 and 6. One test vessel was fabricated from TFP with an Alloy 825 liner. It contained a single butt weld on double seal welded pipe. The second vessel was fabricated from TFP with an Alloy 625 liner and contained two butt welds, one of which joined a double seal weld and one which joined a triple seal welded pipe.

The vessels were modified with pressure fittings so that they could be filled with the test solution and pressurized with a mixed gas containing hydrogen sulfide, carbon dioxide and methane.

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The TFP full scale specimens then were positioned inside of outer containment vessels. The following test conditions were then established inside the specimens: Solution - 20 percent NaCl brine; Gas - 3.5 MPa CO partial pressure, 8.3 MPa H2S partial pressure, balance CH,; To al pressure - 83 MPa; Temperature - 150 2t C. The exposure tests were conducted for a period of 6 months. The pressure in the annulus between the TFP pipe specimen and the containment vessel was monitored during the test for evidence of leakage from the TFP specimen. Gas samples were taken periodically to monitor the constituents in the gas phase.

Following completion of the exposure period the samples were evaluated by the following procedure:

1. End caps removed.

2. I.D. surface of liner tube inspected for corrosion, pits and liner collapse before and after rinsing.

3. The TFP specimens were sectioned longitudinally and further inspected visually and with dye penetrant for evidence of general or localized corrosion, SCC or liner collapse.

4. Metallographic sections were made of selected sites in the weld areas and around any areas of localized corrosion.

Hvdrosen Permeation Studies

The electrochemical permeation technique was used to determine the hydrogen permeation values for various service conditions for hydrogen charging that could take place from either corrosion on the I.D. TFP liner surface or from cathodic protection on the O.D. surface of the TFP. This technique employs two electrochemical cells separated by a thin membrane of the material under consideration [4,5] as shown schematically in Figure 7. The charging side of the membrane, where the hydrogen is deposited, is controlled by a galvanostat. A portion of the hydrogen produced on the charging side of the membrane diffuses through the membrane and exists on the collection side. Here it is measured as a current in the other potentiostatic circuit.

In the present study, three materials were studied for permeability: (1) API X-65 pipeline steel, (2) Alloy 625 and (3) Alloy 825. The membranes used for these experiments were 0.5 mm for the steel and 0.025 mm for the nickel base alloys. The steel membranes were plated with electroless Pd on the measurement side. All membranes were degreased in acetone prior to testing.

Once in the permeation cell, the measurement side was filled with a solution of 0.1 N NaOH and the charging side filled with the test solution. Three test solutions were evaluated: (1) NACE TM0177 solution (5% NaCl, 0.5% acetic acid in distilled water saturated with 100% H,S) at 23 C, 66 C and 88 C, (2) synthetic seawater solution per ASTM 1141 at 23 C, 66 C and 88 C, and (3) 0.51 M sulfuric acid solution containing 0.1 g/l sodium arsenite at 23 C. The cells were purged with nitrogen for one hour; for the TM0177 solution this was followed by purging with hydrogen sulfide.

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The measurement side of the membrane was maintained at 200 mV (SCE) and allowed to stabilize. The charging side of the membrane was varied galvanostatically to three levels of char

P ing current:

(1) 250 pA/cm2, (2) 500 PA/cm', and 1000 PA/cm. The steel membranes were charged for between 2 to 4 hours whereas the nickel base alloy membranes were charged for between 2 to 3 days due to their low hydrogen diffusivity.

RESULTS AND DISCUSSION

Evaluation of CRA Liner with Standardized Tests

The results of the standardized tests for pitting, intergranular (IG) corrosion and fusion line attack of the Alloy 625 liner material are given in Table 2 a-d. These results are presented according to location of the coupons relative to the weldment (LS = liner sample - parent metal, BW = butt weld, and OL = overlay weld - liner seal weld).

ASTM G-48 Ferric Chloride Pittinq Test. As shown in Table 2a, the Alloy 625 specimens did not reveal any evidence of pitting or other form of localized corrosion on the surfaces corresponding to the I.D. surface of the TFP liner. Only slight attack was noted on the coupon surface corresponding to the backside of the TFP overlay seal weld which was initially in contact with the steel outer pipe. This effect results from dilution of the Fe in the steel into the seal weld. However, its effect is limited to the CRA/steel interface which is not exposed to the produced fluids. The data conclusively show that this dilution does not carry forward to the exposed TFP I.D. surface on subsequent weld passes.

ASTM G-28 Ferrous Sulfate/Sulfuric Acid Intersranular (IGL Corrosion Tests. The data in Table 2b, show no evidence of IG attack on any of the parent or weld metal specimens. In general, all welded coupons had greater corrosion rates than parent metal specimens. However, no etching of grain boundaries was detected in any specimens.

HCl Fusion Line Corrosion Tests. The fusion line test consisted of visual observation for localized attack in the unstressed condition and after post test bending around a 0.5 inch mandrel. As in the case of the ASTM G-28 tests, the welded specimens were found to have generally higher corrosion rates than parent metal specimens with very slight fusion line attack on the backside of the specimens. However, as shown in Tables 2c, no evidence of fusion line attack was found in any of the Alloy 625 specimens on the surface corresponding to the TFP liner tube I.D. which would be exposed to the produced fluids.

ASTM A-262C IG Corrosion Tests. As shown in Table 2d, no evidence of IG corrosion was found on any of the specimens evaluated. Furthermore, the results of the tests when evaluated on the basis of weight loss in the fifth exposure period versus the first should not yield more than a twofold increase. In the present study, ratio of corrosion rate in the fifth period versus the first period for all specimens was less than 2, thus indicating acceptable performance.

777

Full Scale TFP Evaluation in Mobile Bav Environment

The full scale TFP pipe vessels were monitored for a period of six months. Only a slight depletion of the gaseous constituents was found to occur during this period. No leakage was noted by monitoring the annular pressure between the TFP specimens and their containment vessels.

After removal from the containment vessels, the TFP full scale specimens were sectioned and inspected. The initial visual and dye penetrant inspections of the two TFP specimens did not reveal any pronounced corrosive attack (localized corrosion or SCC) of the weldments on the CRA liner tubes or liner collapse or blistering. The post-test condition of the Alloy 825 and Alloy 625 liners is shown in Figure 8 a-b.

Small isolated pits adjacent to the overlay region of the Alloy 825 liner material were found situated on surfaces exposed to the liquid phase (Figure 9). These pits were within 5 cm of the Alloy 625 fabrication/overlay weld region. The pits were examined metallographically which revealed that the pitting was not caused by unusual nonmetallic inclusions, anomalous microstructure or surface deformation. As shown in Figure 10, the grain structure of the Alloy 825 liner appeared to be normal in the vicinity of the pits.

From the abovementioned observations, it was concluded that pitting of the Alloy 825 liner in the simulated Mobile Bay production environment is derived from the inherent corrosion resistance of this material relative to the corrosivity of the test environment used in this study. The lack of similar pitting attack in the more corrosion resistant Alloy 625 illustrates this effect. The initiation of pitting in the Alloy 825 could also be influenced by the galvanic interaction between the higher alloyed 625 weld metal and the Alloy 825 liner. In this case, the Alloy 625 would serve as the cathode in the galvanic couple with the Alloy 825 serving as the anode where accelerated anodic attack would be expected.

Hvdroqen Permeation Studies on TFP Liner Collaose

To evaluate the possibilities for TFP liner collapse resulting from hydrogen charging, three cases were evaluated. They consisted of hydrogen charging of (1) steel in a sulfuric acid solution containing sodium arsenite as found in a severe screening test environment for hydrogen charging [l], (2) steel in seawater due to cathodic protection and (3) nickel base alloy liner materials in H,S containing production environments. The data from these experiments is given in Table 3a-b. The values of hydrogen diffusivity and charging efficiency have been used to calculate the hydrogen flux into the steel and the hydrogen pressure build-up rate per methods discussed in the background section of this paper.

It was conservatively assumed that liner collapse will occur when the limiting pressures of 11.5 MPa for an Alloy 625 liner and 8.8 MPa for an Alloy 825 liner are reached. On this bases the useful service life of TFP was estimated.

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Case 1 - Hydroqen Charqinq of Steel in Sulfuric Acid Solutions. From the literature is was noted that in a previous study [l], TFP samples which were hydrogen charged on their 0-D. steel surface in a sulfuric acid solution containing sodium arsenite did not fail by liner collapse after 30 days. By comparison, many other bimetallic tubulars (some with metallurgically bonded liner tubes) suffered collapse of their CPA liner tubes during this same period.

Based on the hydrogen permeation tests conducted in this study, the average life of TFP in the sulfuric acid solution at 23 C with a charging current of 1000 PA/cm2 should be approximately 0.08 years (30 days). This analysis reveals the severe nature of this environment in terms of hydrogen charging. This is due to three major effects: (1) The large amount of hydrogen produced on the charging side of the membrane, (2) the relatively high efficiency of hydrogen absorption in this acidic solution containing sodium arsenite as a hydrogen recombination poison, and (3) the high diffusivity of hydrogen through the steel (relative to stainless alloys). The fact that the previous tests [l] did not result in failure of the TFP samples in the 30 day period is evidence of the conservative nature of the hydrogen collapse calculation.

Case 2 - Hvdroqen Charqinq of Steel in Seawater. In this case, the results pertain to the condition of cathodic protection of the 0-D. steel surface of TFP in seawater over a range of temperatures from 23 C to 88 C. The results are shown in Figure 11 a-b which indicate that the hydrogen diffusivity in steel increases with temperature whereas the permeation efficiency increases from 23 C to 66 C then decreases slightly with increasing temperature to 88 C.

The results of these tests in terms of the service life of Alloy 825/steel TFP yields values of 745 years at 23 C, 64 years at 66 C and 51 years at 88 C. The corresponding useful service lives for Alloy 625/steel TFP will be even greater than determined for Alloy 825/steel TFP due to the fact that the limiting pressure for liner collapse is higher than for Alloy 825 liner. It should be also noted that these results are extremely conservative since the actual percentage of the external surface of an offshore pipeline experiencing cathoding charging is usually very low due to external coatings and calcareous deposits. Assuming the actual area of the pipeline experiencing full cathodic protection would be less than 10 percent, the service lives calculated can be increased by an order of magnitude. Therefore, the shortest estimated service life will be for a hot pipeline and would be in excess of 500 years.

Case 3 - Hydroqen Charqinq of Alloys 825 and 625 in Sour Production Fluids. The *final case under consideration is that of hydrogen charging occurring from corrosion of the Alloy 625 or Alloy 825 liner material in the sour production fluids. In this case, it is assumed that a maximum tolerable corrosion rate on the CPA liner materials is 2 mpy. In order to estimate useful service life in this case, both the hydrogen diffusivities and permeation efficiencies must be known.

Figure 12 a-b show the values of these parameters versus temperature. Alloy 825 appears to have a greater hydrogen permeation efficiency than Alloy 625. However, Alloy 625 shows

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lower hydrogen diffusivities than Alloy 825. The effect of this data on TFP service life is given in Figure 13 for the cases of both Alloy 825 and Alloy 625 liner materials. Over the temperature range 23 C to 88 C, the useful service life of Alloy 825/steel TFP from the standpoint of liner collapse is estimated to range from 289 years at 23 C to 175 years at 88 C. By comparison, corresponding values for Alloy 625/steel TFP are 719 years at 23 C decreasing to 216 years at 88 C.

Overall Suitability of TFP for Mobile Bay Service

The results of the present study have shown that the metallurgical condition of TFP pipeline materials with both Alloy 825 and Alloy 625 is acceptable from the standpoint of microstructure. The standard corrosion tests revealed that both Alloy 825 and Alloy 625 materials can be fabricated to produce suitable liners in TFP.

The inherent corrosion performance of the weldments should be similar to that of the parent metal used for liner construction. The overlay seal welds contain subsurface regions which contain Fe dilution. However, these regions are limited to areas adjacent to the interface between the CPA liner and the steel outer pipe. Furthermore, these regions do not extend to the I.D. surface of the TFP which will be exposed to the sour production environment.

The full scale tests revealed that Alloy 625 lined TFP tubulars can withstand a simulated Mobile Bay production environment containing high levels of hydrogen sulfide and carbon dioxide. No evidence of liner collapse or corrosion damage was noted after a prolonged exposure of 6 months at a temperature of 149 C. However, in the specific environment used for this study, Alloy 825 lined TFP suffered isolated pitting attack of the liner material in the regions exposed to the aqueous phase of the environment. This attack does not appear to be related to metallurgical deficiencies resulting from fabrication. It does appear to relate (1) to the limitations of Alloy 825 relative the corrosivity of the simulated Mobile Bay production environment and (2) to possible galvanic effects between the Alloy 625 weld metal and the Alloy 825 liner material.

Finally, it appears that the overall service life of TFP is not limited in terms of liner .collapse resulting from hydrogen permeation. Very conservative estimates of useful service life Until liner collapse are in all cases greater than 150 years under conditions of liner corrosion and external cathodic protection of the TFP steel outer pipe.

Conclusions

Based on the results of the present study, the following conclusions were made:

1. The TFP (Tight-Fit-Pipe) involving thermal hydraulic fitting, overlay seal welding and fabrication welding exhibits acceptable metallurgical conditions in both Alloy 825 and Alloy 625 liner materials as determined by several standard corrosion tests.

2. The corrosion performance of Alloy 625/steel TFP in a six month full scale exposure to a simulated Mobile Bay

780

production environment showed acceptable corrosion performance with no indication of liner collapse or localized corrosion and SCC on the liner.

3. A similar full scale exposure of Alloy 825/steel TFP revealed no liner collapse but did show evidence of isolated pitting of the liner material near to the Alloy 625 welds. This condition does not appear to be related to metallurgical condition but rather to the limitations of this material in the specific environment chosen for this investigation. Galvanic effects from the Alloy 625 weldments may have had an effect on the initiation of pitting.

4. Hydrogen permeation experiments indicate that TFP liner collapse is not a major limitation of the useful service life of TFP components for Mobile Bay service. Based on calculated hydrogen build-up rates, conservative worst case useful service estimates for all cases are still in excess of 150 years.

Acknowledgements

The authors of this paper wish to thank Kawasaki Heavy Industries, Ltd., of Kobe, Japan, for their support of this research along with the permission to present its findings. Also, acknowledgement is made to Mr. Bradley G. Borgard, Mr. B. Keith Benoit, Mr. Kenneth E. Schindler and other members the technical staff of Cortest Laboratories, Inc. of Houston, Texas, for their efforts in performing this study.

References

1. J-A. Colwell, C-J. Martin and R-D. Mack, "Evaluation of Full Scale Sections of Bimetallic Tubing in Simulated Production Environments", CORROSION/88, Paper No. 57, NACE, Houston, Texas, March (1988).

2. N. Boes and H. Zuchner, "Electrochemical Methods for Studying Diffusion, Permeation and Solubility of Hydrogen in Metals", J of the Less Common Metals, 49, 223 (1976).

3. W-H. Thomason, "Quantitative Measurement of Hydrogen Charging Rates into Steel Surfaces Exposed to Seawater Under Varying Cathodic Protection Levels", Materials Performance, 27, 7, pp 63 - 69, July (1988).

4. M.A.V. Devanathan and Z. Stachurski, "The Adsorption and Diffusion of Electrolytic Hydrogen in Palladium", Proc. Royal Sot. A ., 270, 90, (1962).

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TABLE 1 Evaluation Criteria for Standard Weldment Corrosion Tests

Test Method Evaluation Criteria

ASTM G-48 Coupons are exposed for 72 hours to FeC13 solution and examined for pitting. Test is designed to detect pitting susceptibility of weld metal, liner or overlay. Failure criterion is the observation of pits greater than 10 mil in depth or localized attack of weldment areas.

ASTM G-28

ASTM A-272 Practice C

Fusion Line

Specimen Number

LS LS BW BW OL OL

Test is designed to detect intergranular corrosion caused by the welding process. Coupons are exposed to ~,~~n(SO~)~$i&S04 for .120 -hours.. Weight- loss of the

al exanunatlon diagnose lntergranular attack caused by chromium depletion (carbide formation) during welding. Failure criterion is weight loss of weldment greater than weight loss of base material by a factor of two or more.

Test diagnoses susceptibility to intergranular and/or localized corrosion of susceptible microstructures. Coupons are exposed for five, 48 hour periods to boiling HNO and evaluated for weight loss. Failure criterion 3. -2 weight loss in the fifth period greater than initial weight loss by a factor of two or more. Secondly, weight l'oss of welded coupon should not exceed the weight loss from the base material by a factor of two or more.

Boiling HCl is used to produce a high rate of general corrosion. Fusion lines are visually examined to detect selective dissolution. Some coupons are bent over a mandrel to detect hair line attack at the fusion interface. Failure criterion is selective (greater than 10 mil penetration) fusion line dissolution.

TABLE 2A ASTM G-48 Test Data

Material Number

625-2A 625-21 625-3A 625-3B 625-3B 625-3~

Description

No pitting or localized attack No pitting or localized attack No pitting or localized attack No pitting or localized attack No pitting or localized attack Backside FL attack (very slight) (1)

(1) Fusion line between the line and overlay pass on side adjacent to steel.

TABLE 2B ASTM G-28 Test Data

Specimen Material Number Number Description

LS 625-2~ No intergranular attack LS 625-2~ No intergranular attack BW 625-3~ No intergranular attack BW 625-3B No intergranular attack OL 625-3~ OL

No intergranular attack (1) 625-3~ No intergranular attack

(1) Slight corrosion on side adjacent to steel at FL of overlay and liner.

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Specimen Number

LS LS BW BW OL

OL

Exposure

A B C

A B C

A B C

A B C

Material Number

625-214 625-2B 625-3A 625-3B 625-3A

As Exposed After

General Corrosion ----- General Corrosion _---- General Corrosion No FL Attack General Corrosion No FL Attack Sliuht Backside FL Attack (1) No FL Attack

625-3B General Corrosion

Slight Backside FL Attack (1) No FL Attack General Corrosion

TABLE 20 ASTM A-262 Practice C

Specimen Number Material Number

BW 625-3A BW 625-3A BW 625-3A

BW BW BW

625-3B 625-3B 625-3B

OL OL OL

OL OL OL

625-3A 625-3A 625-3A

625-3B 625-3B 625-3B

TABLE 2C Fusion Line Test Data

Description

Ratio Description mpy S/mpy 1

No localized pitting 1.7 No localized pitting No localized pitting

No localized pitting No localized pitting No localized pitting

1.6

No localized pitting No localized pitting No localized pitting

1.8

No localized pitting No localized pitting No localized pitting

i-4

TABLE 3A Alloy Permeation Calculations

PC0 Temp. Uxsrgmg Chargmg Stew state Alpha (dlffuswtvJ) (thicknessi liner lpa moles nf psi from nf years 4LLOY Solution mA/cmZ lco (mA/cm2! ICQ/lC (cmZ/sec) 1 (cm) (mA/cm2) :nto liner Into liner failure

825 75 F NACE 0.125 2.99E-03 0.0239 I .39E-08 2.54E-03 l.OlE-36 5.24E- IS I .88E-07 215 825 75F HZS 0.250 3.53E-03 0.0141 7.62E-09 2.54E-03 5.98E-07 3. IOE- IS I .1 IE-07 363 825 75 F sat. 0.500 2.54E-03

Average Values= 0.019 I OSE-08 8.05E-07 4.17E- I5 1.50E-i)7 289

825 150 F NACE 0.125 2.96E-03 0.0237 7.10E-08 2.54E-03 I .OOE-06 5.20E- 15 2. IZE-07 190 825 150F HZS 0.250 522E-03 0.0209 2.54E-03 8.83E-07 4.58E- I5 I .87E-07 2 16 825 ISOF sat. 0.500 7 17E-03 0.0143 2.54E-03 6.07E-07 3.14E- 15 I .28E-07 315

Average Values= 0.0 196 8.3lE-07 4.31E-15 1.76E-07 241

825 190 F NACE 0.125 3.89E-03 0.031 1 2.67E-57 2.54E-03 I .32E-06 6.82E- 15 2.97E-07 136 825 19OF H2S 0.250 5.78E-03 0.0231 2.54E-03 9.79E-07 5.08E- 15 2.21 E-07 183 825 19OF sat. 0.500 I 03E-02 0.0206 254E-03 B.flE-07 4.5lE-15 196E-07 206

Average Values= 0.0273 I 06E-06 5.47E- I5 2.38E-07 175

625 75F MACE 0.125 7.62E-03 625 75F H2S 0.250 7.62E-03 625 75 F sat. 0.500 I 55E-03 0.0031 6.18E-38 7 6X-03 3.95E-07 2.05E- 15 7 35E-08 719

Average Values= 3 95E-07 2.05E- i5 ?.35E-08 719

625 150 F NACE 0.125 1 19E-03 0.0095 2.09E-07 7 62E-03 I .2 1 E-06 6.27E- IS 2.56E-07 206 625 ISOF H2S 0.250 1.58E-03 0.0063 2.09E-07 7 62E-03 8.03E-07 4.16E- 15 1.70E-07 311 625 15OF sat. 0.500 1 97E-03 0.0039 2.09E-07 7 62E-03 S.OlE-07 2.60E- 15 I .06E-07 499

Average Values= 0.0066 8.38E-07 4.34E- i 5 I .77E-07 339

625 190 F MACE 0.125 I .28E-03 0.0102 I .87E-07 7.62E-03 1.30E-06 6.73E- I5 2.93E-07 181 625 19OF H2S 0.250 2.46E-03 0.0099 4.00E-07 7.62E-03 I .25E-06 6.48E- 15 2.82E-07 187 625 19OF sat. 0.500 3.29E-03 0.0066 4 OOE-07 7.62E-03 8.36E-07 4.33E- 15 1 88E-07 280

Aver age Values= 0.0089 3.29E-07 I 13E-06 S.8SE- 15 Z.SSE-07 216

783

TABLE 38 Steel Permeation Celculatlons

mam. Temp. Charging Chargfng Stew state Alpha (diffusfvitv) (thmknass) liner lee inoles nf psi irom nf yaars to Type F solution mA/cm2 100 (mA/cmZ) IWIc D I (cm) (mA/cm2) lnt0 liner into liner failure

steel 75 F sulfuric 0.125 I .90E-02 0.1524 8.17E-06 4.83E-02 7.93E-03 4. I 1 E- l I 1.48E-03 0.03 steal 75F acfd 0.250 l.l7E-02 0.0469 1.89E-05 4.836-02 2.44E-03 I .26E- I I 4.54E-04 0.09 steel 75 F arsenate 0.500 I .90E-02 0.038 9.60E-06 4 83E-02 I 98E-03 l.O2E-1 I 3.68E-04 0. I I

Aver- values = 0.0791 I .22 I E-05 4.12E-53 213E- I I 7.67E-04 0.08

steel 75 F Synthettc 0.125 2.7 lE-04 0.0022 2.OSE-05 4.83E-02 564E-07 2.92E- 15 1.05E-07 305 stwl 75F Seawater 0.250 3.18E-04 0.0013 Z.OSE-05 4.83E-02 3.32E-07 I .72E- 15 6.185-08 655 steel 75 F sclution 0.500 3.48E-04 0.0007 3.43E-05 4.83E-02 1.81E-07 9 40E- 16 3.38E-08 I 197

Average values = 0.0014 2.513E-05 3,59E-07 I.&E- 15 6.69E-08 745

steel 150 F Svnthetfc 0. I25 l.l2E-03 0.0089 3.12E-05 3.30E-02 I .59E-06 8.26E- I5 3.37E-07 120 steel I50 F ,Seawater 0.250 9.44E-03 0.0378 7.69E-05 3.3DE-02 &73E-06 3.49E-14 1.42E-06 28 steel 150 F solution 0.500 I 22E-02 0.0244 6.42E-05 3 30E-02 4.35E-06 2.26E-14 9.2lE-07 44

Average values = 0.0237 5.743E-05 4,22E-06 2.19E- I4 &93E-07 64

steel I90 F Svnthetlc 0.125 ?.70E-03 0.0216 7.44E-05 4.95E-02 5.78E-36 2.99E-14 1.30E-06 31 steel I90 F r%awater 0.250 3.52E-03 0.0141 7.39E-05 4.95E-02 3.76E-06 1 95E-14 8.49E-07 48 steel 190 F solution 0.500 4 43E-03 0.0089 9.50E-05 4.95E-02 :.37E-06 I 23E- 14 5.34E-07 76

Average values = 0.0149 8.i 1 IE-05 3,97E-06 2.06E- I4 8.95E-07 5 I

\t

a Dm Outer Pipe

Cooled

FIGURE 1 - THERMOHYDRAULIC FIT METHOD

784

(01 As tight -fitted

tbJ Machining

Weld

An Example of Dimension ( TFP for 14AM 1 -__-- ----.-

_ DP.A.?-~? !!‘I’! _ -IP: 9, 3 !Ilr!l- _ t : -- 3. Q !v _ -d-;-3.5 mm- _ -.L.3.0 rwn- r - Do:! 5.9 !nrn-

tdl Edge

FIGURE 2

preparation

- SEAL WELD IREATHEHT OF 1FP ENDS

Seal Weld Fusion Line

’ Liner

Figure 4 - Coupon Locations in Pipe Mall for 625-3

Weld Join1

Detail of Weld Joint

Seal Weld ( l-625 : z-larerr )

Weld Join1 (3L) Weld Joint (ZL)

PIUP, l-625 )

Detail of Weld Joint (21)

Outer Pipe -

( l-625 : z-lsrers ) T ”

Detail of Weld Joint (31)

Outer Pipe - ( API X65 )

Bull Weld /’ l-625 1

Seal Weld

t .” p‘ ? n

‘,L 0,

20

rr

II’

\,

& _ ‘*,’

,,, ,,>~ ,‘,#m.d

-a :.

* ,- 1. _ )L_(, -/ _( :

(, *l ,,,, _ ,_,,: ; ‘,,,, i

Figure 8A - Post-test ID surface conditions of Alloy 825/Steel TFP specimen.

Figure 8B - Post-test ID surface conditions of Alloy 625/Steel TFP specimen.

788

iI

Y

I

il ‘ I * la c* 1 :,

Figure 9 - Close-up view of isolated pits Alloy 825 liner after cleaning.

in

789

790

800

700

600

500

400

300

200

100 60 80 100 120 140 160 180 200

Temper8mre F

‘Figure 13

025

- 625

I

791