Electromagnetic Non-Destructive Evaluation II,
Proceedings of the 3rd International Workshop on E'NDE,
Reggio Calabria, Italy, September 1997.
ISBN: 90 5199 375 7 - published by IOS Press
|TABLE OF CONTENTS|
Successful remote-field eddy current examination of high-pressure feedwater heaters in the power industry requires a comprehensive inspection program backed by qualified NDE techniques and operators. If applied effectively, NDE can extend availability of operating heat exchangers, reduce maintenance costs by minimizing tube leaks, and offer timely heat exchanger replacement by providing reliable analysis of successive inspection data.
Since 1991, the EPRI NDE Center has acquired, characterized, and assembled a compendium of realistic BOP heat exchanger mockups. The mockups containing both corrosion and mechanical based damage forms are used to support EPRI member utilities in performance demonstration tests, training, and technology transfer programs. This paper details more recent remote-field eddy current applications to carbon steel tubing containing actual field-induced OD-initiating flaws.
In the last ten years, this technique has emerged from its application in the oil country tubular goods to inspection of smaller diameter heat exchanger tubing, . Although it can be used effectively for both non-magnetic and ferromagnetic tubing, its application is confined mainly to ferromagnetic tubing, such as carbon steel and 439 ferritic stainless steel. This technique, unlike the flux leakage or mag-biased eddy current technique, requires no external biasing field to evaluate the integrity of ferromagnetic tubing.
|Fig 1b: Relative Signal Amplitude and Phase Orientation at Three Different Zones|
Fig 1a: Receiver Coil(s) Located in the Remote-Field Zone
As shown in Figure 1, the primary field has effectively penetrated the tube wall from ID to OD, traveled along the OD tube surface, and reentered the tube wall to be detected by the ID receiver coil(s) located approximately 3 to 4 tube diameters away from the exciting coil. This region, identified as a remote-field zone, is physically separated such that no direct coupling on the ID surface takes place from the exciter coil. This is due to the field attenuation and is caused by induced eddy currents generating a secondary field that opposes the primary field, thus effectively shielding and preventing any direct coupling with the receiver coil(s). Because of the double wall entry by the primary field, this represents a volumetric tube wall inspection technique.
Among the electromagnetic NDE techniques evaluated in the past (see Table 1) for SA 556-C2 carbon steel tubing, 3/4-inch (19.1 mm) x 0.109-inch (2.77 mm) nominal wall, the remote field eddy current technique provided the best overall capability to characterize OD-initiating flaws. Damage mechanisms consisted mainly of steam erosion, steam cutting next to tube support plates, and through-wall ruptures.
As in the traditional eddy current impedance plane analysis for OD flaws, the phase angle analysis by remote-field eddy current provided the best overall flaw depth sizing results. This was accomplished by analyzing the complex voltage plane trajectory (Lissajous figure) of the real (in-phase) and imaginary (quadrature) components of the induced receiver voltage, .
It was noted, however, that the presence of tube support plates shifted the analysis results from being conservative to nonconservative regardless of the techniques used, for example, flaw depths were underestimated in the presence of tube support plates.
|Mag-Bias (1.2 kHz)||Flux Leakage (DC)||Remote Field (250 Hz)|
Unlike the older remote-field inspection probe consisting of an exciter and single receiver coil, the latest probe incorporates a dual-exciter and differential receiver coils as shown in Figure 2a. Two similar size exciting coils are placed three tube diameters away from the receiver coils and each exciting coil is driven in opposition by the low frequency AC current. This probe equipped with a 50 foot long cable length was connected to a dual-frequency MIZ-40RFT instrument shown in Figure 2b.
|Fig 2b: Dual-Frequency MIZ-40RFT Instrument|
Fig 2a: A Dual-Exciter with Differential Receiver Coils
The following instrument specifications were provided:
Based on the damage forms known to exist in the testing mockup, three different calibration standards were used to analyze the acquired data. These standards consisted of
The actual data acquisition was performed using a 0.450 inch (11.43 mm) diameter probe. This represented a fill-factor of around 72% for testing the high-pressure feedwater tubing mockup. Since the inspection vendor had no access to the actual tube with a similar size wall thickness, a set of phase angle-to-flaw depth calibration curves were established at a given frequency using the above probe on reference standards containing less than 0.109-inch nominal wall thickness.
Separate calibration runs were conducted with and without support rings over individual flaws. Thus, two sets of calibration curves were established: one for mid-span tube indications and another for flaws in the vicinity of tube support plates. Figure 3 shows the calibration curves for evaluating flaws located next to the tube support plates.
|Fig 3: Phase Angle-to-Flaw Depth Calibration Curves with Tube Support Rings Over Individual Flaws from Three Different Standards|
Fig 4: Normal Tube Support Plate Responses in Differential and Absolute Modes
Fig 5: Remote-Field Signal Responses in Free Tube Span
Fig 6: Non-Through-Wall Indications at Tube Support Plate #3
Fig 7: Through-Wall Indications at Tube Support Plate #3
Fig 8: Flaw Sizing Results for 0.109" Wall Carbon Steel Tubing
Before performing the actual examination, an optimum frequency was selected by comparing the pre-prepared calibration curve signal responses with those signals obtained from the actual mockup. This was accomplished by comparing common signal responses representing a through-wall hole and a support plate. After the comparison, an operating frequency of 380 Hz was eventually selected. At this selected frequency, the tube support plate response was set to 210° @ 5 volts. Figure 4 shows the typical support plate responses obtained from the mockup in both differential and absolute modes.
After normalizing the tube support plate responses to selected gain and phase angle rotations, the through wall responses were obtained at 90° and 50°, respectively, in differential and absolute modes. As shown in Figure 5, a 20% groove response is rotated clock-wise to around 162° and 127°, respectively. This resulted in more than 70° of phase spread over the 20-100% through-wall range in the free tube span regions
During the course of the examination, it was noted that tube support plate responses were not maintained constant in terms of their amplitude and phase angle orientations. This variation was attributed to permeability variations along the length of tubing and between the individual tubes. To minimize the effect and to provide more consistent analysis results, it was decided to normalize a clean support plate response found in each tube before obtaining any flaw response measurements. This was accomplished by manually adjusting the clean support plate response to 210° at 5 volts, peak-to-peak.
Once the flaws were detected and measured, the next step was to select an appropriate calibration curve for estimating flaw depths. For this mockup, the primary damage forms consisted of steam erosion and through-wall holes next to the tube support plates. Consequently, a set of calibration curves with support plates over individual flaws was selected to assess the flaw depth. Measured signal amplitude and phase angles resulted in the use of the 360° erosion curve for non-through-wall indications. For through-wall indications, the 180° erosion curves was used (see Figure 3).
Figures 6 and 7 show examples of non-through-wall and through-wall indications, respectively. Dual exciter coils allowed an easier interpretation of flaw signals since they were very similar to conventional eddy current impedance plane trajectories. The only difference is the presence of additional flaw signals that are present due to the exciter coils traversing over the flaws. Data analysis was performed primarily by equating the differential phase angles to flaw depths using the 360° erosion curve. Similarly, the flaw confirmation was accomplished by reviewing the absolute signal responses.
The overall performance demonstration involved both flaw detection and sizing tests. Only those indications with flaw depths exceeding 20% wall losses were included in the evaluation. All of the flaws exceeding 20% wall losses were detected successfully with no false calls from those control tubes containing no flaws.
To obtain statistically meaningful sizing data from the mockup containing 36 tubes, the data analysis was conducted by two different operators. Figure 8 shows graphically the combined analysis results of two data analysts.
An overall improvement in flaw sizing was obtained by using the dual exciter remote-field probe combined with the use of a calibration curve that takes into consideration the presence of tube support plates. This improvement is evident by comparing the present and past statistical values that were used to assess the sizing accuracy: correlation coefficient of 94% versus 87% and RMS error of 9% versus 20%. The overall slope of the correlation curve remained about the same (0.89 versus 0.91).
It is possible, based on the phase angle analysis, to discriminate ID- from OD-initiating flaws by reviewing their respective phase angle orientations. Also, all flaw-like signals initiate in counter clock-wise directions, while OD artifacts such as tube support plates form in clock-wise directions (see Figures 4 and 6).
Because of differences in the calibration curves for varying damage forms, more accurate flaw sizing results are possible by using appropriate standards that emulate flaw types to be detected and characterized.