Aug 22, 2014 - coupling-based telemetry system equipped with two electronic gauges which provide ... dition, it makes it possible to supply power to downhole apparatus ... the coupling factor M decreases with the number of turns n2, which.
Coreless Electromagnetic Coupling-Based Drillstem Telemetry Using Dual Electronic Gauges Tianhuai Ding and Li Cheng, Tsinghua U.
Summary A coreless electromagnetic coupling-based telemetry system for well-test data involving a full-bore drillstem test apparatus has been developed. This system allows the retrieval of formation pressures and temperatures above and below a tester valve by means of a wireline-conveyed electronic gauge and a permanently installed gauge. A wireline-deployed coreless electromagnetic proximity sonde is put into the well to transfer the data using time-division frequency transmission technology in the “Synchronization-Pressure-Low-Temperature-Low” mode on a single-core wired-armored cable. The coded pulse data are modulated on DC power signals. The surface processor uses a double-counter synchronization method to measure signal frequencies to eliminate ±1 most significant bit (MSB) error. Then a simplified algorithm is proposed, based on least-squares curve fitting and dimensionreduction methods, for computing actual temperatures and pressures for analysis of well-testing data. The proposed system is calibrated for temperatures up to 125°C and pressures up to 70 MPa with a pressure accuracy of 0.1% full scale (FS) and temperature accuracy of ±0.5°C. This approach combines the advantages of drillstem testing and wireline formation testing and has been successfully applied to onshore well testing. Introduction In oil and gas well testing, it is important to obtain accurate realtime pressures and temperatures from the bottom of a well. The need for surface-to-downhole communication in this context has long been recognized, and many techniques have been proposed, including wireline telemetry (Whittle et al. 2003), permanently installed downhole monitoring systems (Veneruso et al. 2000), and wireless telemetry (“Electromagnetic” 200l; Tochikawa et al. 1996). The wireline telemetry system provides higher data rates than other currently used methods. However, it requires that the sensitive measurement gauges endure long-term exposure to an extremely hostile environment. The permanently installed downhole monitoring system records data in a module that is retrieved after the tubing is lifted from the hole. However, this method makes it difficult to collect real-time data. Wireless telemetry involves either extremely low-frequency electromagnetic (EM) wave transmission through the formation or acoustic transmission through the drillstring. The propagation of EM waves is characterized by an increase in attenuation with increases in distance, data rate, and ground conductivity (Trofimenkoff et al. 2000). Therefore, current EM telemetry systems are practical only for land operations where the resistivities are greater than 1 ohm-m and the target formation is shallow (Soulier 2003). Though Halliburton has recently developed an acoustic telemetry data acquisition system (Harper et al. 2003), it is limited by its high energy consumption and the complicated problems caused by the lengthwise variations of drillstems. This paper describes a coreless EM coupling-based telemetry system equipped with two electronic gauges which provide users with access to real-time downhole data above and below the tester valve during drillstem testing operations.
Copyright © 2007 Society of Petroleum Engineers Original SPE manuscript received for review 2 August 2005. Revised manuscript received 29 March 2006. Paper (SPE 99365) peer approved 29 March 2006.
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Telemetry System Structure. The system is used primarily in conjunction with 5-in. OD drillstem testing tools (Fig. 1). In contrast to conventional drillstem testing systems, it uses both a wireline-conveyed gauge and a permanently installed gauge. The permanently installed gauge is lowered into the well with the tubing. The wirelineconveyed gauge, connected to a sonde, is run at the top of the tester valve in the hole. The valve is then opened to flow the well and later closed to build up the well. The gauges provide accurate downhole pressure and temperature (P/T) measurements, both in memory and in surface readout mode. Using different DC power signals (27 V or 16 V) on the electric cable, a specific gauge is available to record downhole data during the drawdowns that occur during well flow or buildup. The gauge in the tubing is coupled with a large coreless coil which transmits the data for reception by another small coreless coil positioned in the sonde. Then pulse data coupled with DC power are transferred to the surface by the cable. A simplified fitting algorithm in the surface processor is used to compute actual pressures and temperatures (P/Ts) for subsequent formation analysis and permeability prediction. Advantages. The system offers the advantages of both drillstem testing and wireline formation testing. The wireline-deployed sonde can be lowered into the hole at any appropriate time, which enables real-time measurement and reduces the loss of data and time resulting from the failure of a permanently installed gauge. By taking readings at several test points, it can provide reservoir delineation by means of pressure-gradient plots. The permanently installed gauge can obtain integrated history data and compensate for the distortion caused by a substantial volume of mud filtrate invading the formation after the zones of interest have been drilled. The system eliminates the shortcomings of high attenuation and shallow testing depth faced by current wireless telemetry. In addition, it makes it possible to supply power to downhole apparatus during well-testing operations. Another advantage of the system is that it implements powersaving features by means of a noncontact inductive proximity switch. The switch consists of a controller in the sonde and an actuator connected to the battery packs in the tubing. A transmitter coil in the controller generates an EM field. When the sonde approaches downhole tools, the transmitter coil is coupled with the receiver coil in the actuator. Then the inductive electromotive force can activate the switch and perform on-off control of the downhole battery packs. The effective working distance is ±80 mm. The power supply for the switch comes from inductive coupling instead of the battery packs, which extends battery life. Finally, the system overcomes the problems of the bad contacts and seals that downhole contact signal transmission methods commonly encounter. Coreless coils, by eliminating cores of special materials (Veneruso 1990), further simplify the structure of the system, as shown in Fig. 2. Electrically nonconductive reinforced plastic, which is coaxially arranged around the coil assemblies, physically protects the coils. The two-injection sealed technology prevents caustic liquids from penetrating the coils. System Design Fig. 3 gives a brief overview of the telemetry system. First, the coreless EM coupling is described, and then the wireline remote transmission and P/T solution algorithms are presented. February 2007 SPE Production & Operations
Fig. 1—Standard configuration of telemetry system.
Coreless Electromagnetic Coupling. EM coupling is governed by Faraday’s law of induction and satisfies
内 Edl = − ddt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (1) B
An alternating-current magnetic-dipole field resulting from EM coupling in the radio frequency band is near field with little or no electric field. The magnetic field is nonpropagating and is primarily a diffusion field in a conductive medium. Coil parameters are optimized to reduce stray capacitance and improve coupling efficiency according to the magnetic moment m through the central axis of a magnetic dipole. m can be calculated as m = EnIA, . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (2) where E is the effective permeability of the cored coil, n is the number of turns, I is the current, and A is the cross-sectional area of the coil. m can be increased by increasing any of E, n, I, and A. However, there are tradeoffs. The use of a permeable core material comes at the cost of core power losses, and the material itself becomes scorched in the downhole high-temperature environment, which makes coreless EM coupling technology superior for this application (Ding et al. 2000). Increasing n can result in greater core losses and a higher rate of copper loss. Increasing I can lead to increased power losses in the copper resistance element, as well as core losses. Increasing A is an effective approach to increase moment, particular with coreless coils, because a doubling of coil diameter increases the moment by a factor of four. However, the feasible range of these parameters is limited by tubing dimensions. Therefore, the diameter d2 and the length l2 of the large coil coupled with the permanently installed gauge can be determined as 66 mm and 20 mm, respectively. The selfinductance L2 of the coil can be calculated as L2 =
E 2 n d ␣ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (3) 4 2 2 2
Fig. 2—Standard structure of coreless coils. February 2007 SPE Production & Operations
Fig. 3—The structure diagram of the system.
Assuming the lengths of both the large l and the small coils are equal to l, from M=
2n2 d22 2␣1n1l2
冉冑 冉 冊 冊 1+4
l d1
2
− 1 , . . . . . . . . . . . . . . . . . . . (4)
the coupling factor M decreases with the number of turns n2, which degrades EM coupling. On the basis of experiments, a value of 30 can be determined for n2. The diameter d1, length l1, and number of turns n1 of the small coil in the sonde are 28 mm, 20 mm, and 120 mm, respectively. The inner coil assembly in the sonde is coupled with the coil assembly that is coaxially arranged around the axial passage in the tubing, with the three-point capacitance oscillator circuit based on on-off keying (OOK) modulation. Wireline Remote Transmission. The temperature gradients between topside and downhole locations are very large and affect transmission parameters and cable performance. Conventional methods of cable matching, such as using a purely resistive source and load impedance equal to the cable characteristic impedance, do not necessarily deliver optimal performance when using long cables for transmission of low to medium frequencies. The characteristic impedance of the cable is a function of frequency and temperature. Qureshi et al. (1999) have developed a mathematical model based on the primary constants of the transmission line at any given temperature. The model is used to predict the power output when a 1-km cable is subjected to 100°C. The simulation results indicate that when the cable is dynamically matched, the 129
Fig. 4—Data transmission mode.
power output increases by approximately 0.3% (Qureshi et al. 1999), which seems to represent no significant advantage. To implement P/T remote transmission in cable pairs, the system uses time-division frequency transmission technology in the “Synchronization-Pressure-Low-Temperature-Low” (“S-P-L-TL”) mode (Fig. 4). S denotes a 250 Hz pulse used for data synchronization, P and T represent the P/T frequencies, and L is the low-level separator symbol for the P/Ts. The sonde sends a packet to the surface every second. The time-span ratio among the parameters in one transmission period is 5:1:3:1. The surface processor uses double pulse counters to measure the data frequency fx and the clock signal frequency f0 during the fixed sampling time, as shown in Fig. 5. The capture register, when enabled, stores the count in the free-running timer at the rising edge of an input pulse signal. This rising edge also serves as the clock for the pulse counter. The system starts the prescribed gate and the synchronizing gate. The double counters run until the next rising edge of an input signal after the prescribed time. After that, the capture function is again enabled. Then, by using the pulse count of the input signal Nx and the clock signal N0 during the prescribed time and the clock frequency, the signal frequency can be defined as fx =
Nx f . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (5) N0 0
The single-core wired-armored cable is also used for communicating electricity to the sonde and the wireline-conveyed gauge. A regulated voltage is transmitted through the 1-wire from the well surface equipment to the downhole tools. The P/T signals are sent to the well surface by modulating DC power signals on the 1-wire. Especially during work at low levels, the surface processor can send additional commands or introductions downward, enabling communication in both directions. P/T Fitting Algorithm. The actual P/T values in well testing are of the utmost significance (Van Riet et al. 2004). A simplified fitting algorithm based on the sensor characteristics of the system is presented to compute these data for subsequent surface well analysis. The raw data collected by the P/T resistor and sputtering thinfilm resistive pressure sensors in the electronic gauge are converted to frequencies by means of a voltage-to-frequency converter for reliable remote transmission. After the P/T frequencies are obtained, the actual P/Ts can be determined by:
再
Tជ = FT共fជT兲
, . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (6) Pជ = FP共fជT, fជP兲
where FP and FT are inverse sensor transfer functions, fជP and fជT are acquired P/T frequency vectors, and Pជ and Tជ are actual P/T vectors. Temperature Solution. The relationship between the P/T thermal resistance R and the temperature T is RT = R0共1 + r1T + r2T 2 + r3T 3 + · · · + rNT N兲, . . . . . . . . . . . . . (7)
Fig. 5—Pulse frequency measurement.
Assuming that the frequency variation is linear with the resistance variation as the temperature changes, the relationship between the temperature T and the corresponding frequency f can also be expressed in a polynomial form similar to Eq. 8, namely T = c0 + c1 f + c2 f 2 + c3 f 3 + e, . . . . . . . . . . . . . . . . . . . . . . . . . . (9) where e is the fitting error and cj(j⳱0, 1, 2, 3) are temperature/ frequency coefficients. Eq. 9 can be rewritten in vector form as Tជ = c0 fជ0T + c1 fជT + c2fជ2T + c3 fជ3T + eជ, . . . . . . . . . . . . . . . . . . . . . . . (10) where Tជ ⳱[T0, T1, . . . , TN−1]T is the temperature vector, fជT⳱[ fT0, fT1, . . . , fTN−1]T is the frequency vector, eជ⳱[e0, e1, . . . , eN−1]T is the fitting error vector, and fជ0T, fជ2T, and fជ3T are defined as fជ0T⳱[ f T0 0, f T0 1, . . . , f T0 N−1]T⳱[1, 1, . . . , 1]T, fជ2T⳱[ f T2 0, f T2 1, . . . , f T2 N−1]T, and fជ3T⳱[ f T3 0, f T3 1, . . . , f T3 N−1]T, whose components are respectively the fT. 0th, second, and third powers of the components of vector ជ Then the fitted polynomial expression for the actual temperatures is given by
共fជT兲 = c0fជ0T + c1fជT + c2fជ2T + c3fជ3T. . . . . . . . . . . . . . . . . . . . . . . . (11) Assuming N(Nⱖ4) temperatures in the vector Tជ ⳱[T0, T1, . . . , TN−1]T are known, the 2-norm least-squares goodness-of-fit criterion implies that Y共c0, c1, c2, c3兲 = 㛳Tជ − 共 fជT兲㛳2 =
N−1
兺 关T − 共f 兲兴 . 2
i
i=0
Ti
. . . . . . . . (12)
The condition for the minimum fitting error eជ⳱[e0, e1, . . . , eN−1]T to exist is either ⭸Y = 0. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (13) ⭸cj or ជ兲, 共Tជ ·  ជ兲, 共Tជ ·  ជ兲兴T, . . . . . . . . . . . (14) ជ = 关共Tជ ·  ជ兲, 共Tជ ·  G*C 0 1 2 3
where G =
冤
ជ· ជ兲 共 0 0
ជ· ជ兲 共 1 0
ជ· ជ兲 共 2 0
ជ· ជ兲 共 3 0
ជ· ជ兲 共 0 1
ជ· ជ兲 共 1 1
ជ· ជ兲 共 2 1
ជ· ជ兲 共 3 1
ជ· ជ兲 共 0 2
ជ· ជ兲 共 1 2
ជ· ជ兲 共 2 2
ជ· ជ兲 共 3 2
ជ· ជ兲 共 0 3
ជ· ជ兲 共 1 3
ជ· ជ兲 共 2 3
ជ· ជ兲 共 3 3
冥
,
where RT is the resistance at temperature T°C, R0 is the resistance at temperature 0°C, and the ri(i⳱1, 2, . . . , N) are resistance/ temperature coefficients. In the downhole temperature range between 0°C and 150°C, Eq. 7 can be simplified as
ជ ⳱[c , c , c , c ]T, and the operator (aជ ·bជ ) ជ= fជj 共 j⳱0, 1, 2, 3兲, C  j T 0 1 2 3 represents the dot product of vectors aជ and bជ . f 3T are linearly independent, Because the vectors ជ f 0T, fជT, fជ2T, and ជ the determinant of the matrix G must be nonzero. Then the temជ is perature/frequency coefficient vector C
RT = R0共1 + r1T + r2T 2 + r3T 3兲. . . . . . . . . . . . . . . . . . . . . . . . . . (8)
ជ = G −1*关共Tជ ·  ជ兲, 共Tជ ·  ជ兲, 共Tជ ·  ជ兲, 共Tជ ·  ជ兲兴T. . . . . . . . . . . . (15) C 0 1 2 3
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Fig. 7—The effects on receiver coil voltage.
Fig. 6—The effects on transmitter coil voltage.
ជ and the acquired Substituting the fitting coefficient vector C frequency vector fជT into Eq. 11, we can obtain the actual fitted ជ(T*, T*, . . . , T* ). temperature vector T* 0 1 N−1 Pressure Solution. The pressure is a bivariate function of fជT and fជP. Although a 2D cubic spline fitting method can solve the curved surface representing the pressure, the computations will consume large amounts of memory in the surface processor. We present a dimension-reduction fitting method for the computation of the actual pressures by using an ordinary cubic spline fitting function at grid boundaries. The fitted pressure solution at a fixed temperature T is similar to the temperature solution and is given by fជP = k0 fជ0P + k1 fជP + k2 fជ2P + k3 fជ3P, . . . . . . . . . . . . . . . . . . . . . . . . (16) where kជ(k0, k1, k2, k3) is the pressure/frequency coefficient vector, Pជ is the fitted pressure vector, fជP⳱[fp0, fp1, . . . , fpN−1]T is the pressure frequency vector, and ជ f 0P, fជ2P, and fជ3P are given by fជ0P=关 f 0p0, 0 0 T f p1, . . . , f pN−1兴 ⳱[1, 1, . . . , 1]T, fជ2P⳱[f p20, f 2p1, . . . , f p2N−1]T, and fជ3P⳱关f 3p0, f 3p1, . . . , f p3N−1兴T, whose components are respectively the 0th, second, and third powers of the components of vector ជ fP. ជ, P ជ, . . . Then the corresponding fitted pressure vectors ជ P0, P 1 2 ជ at N different temperatures T , T , T , . . . T P N−1 0 1 2 N−1 satisfy
冦
ជ = k fជ0 + k fជ + k fជ2 + k fជ3 共T = T 兲 P 0 00 P 01 P 02 P 03 P 0 ជ = k fជ0 + k fជ + k fជ2 + k fជ3 共T = T 兲 P 1 10 P 11 P 22 P 23 P 1
where ជj(j,0, j,1, j,2, j,3) ( j⳱0, 1, 2, 3) denotes the fitting coefficient vector for the pressures at different temperatures. Therefore, the column coefficient vector kជj( j⳱0, 1, 2, 3) can be ជ determined by substituting T * and the vector ជj(j,0, j,1, j,2, j,3)( j⳱0, 1, 2, 3) into Eq. 18. Further substitutions of the acquired fជP and the corresponding components of the coefficient matrix K共k0, kជ1, kជ2, kជ3) into Eq. 17 can yield the actual fitted pressure ជ(p*, p*, . . . , p* ) at the temperature vector ជ T*(T* vector P* 0 1 N−1 0, T* , . . . , T* 1 N−1). Experiments and Applications Coreless EM Coupling. Because the inductive coils are surrounded by thick-walled drillpipes, a significant amount of electrical energy that might otherwise be transferred through these coils will instead be dissipated into the electrically conductive pipe. EM coupling experiments were performed to analyze these coupling effects by using 1-meter-long stainless steel pipe and coreless coils with an inner radius of 110 mm and an outer radius of 122 mm (Ding and Li 2004). Figs. 6 and 7 show the effects of temperature on transmitter coil voltage and inductive coil voltage. Fig. 8 shows the effects of temperature on coupling distance. The shielding effects of the cylindrical tube on self-inductance and coupling distance are summarized in Table 1. The results suggest that temperature has little effect on inductive coil voltage and coupling distance, with or without tube shielding. Secondly, the metal shielding creates an increase in self-inductance, but attenuates the coupling distance from 40 to 25 mm. However, the system can position the sonde downhole for reliable EM coupling using a latch device.
...... ជ3 ជ0 ជ2 ជ ជ=k P N−1 N−1,0 f P + kN−1,1 fP + kN−1,2 f P + kN−1,3 f P共T = TN−1 兲 . . . . . . . . . . . . . . . . . . . . . . . . . . (17)
where
K = 关kជ0
kជ1
kជ2
kជ3兴 =
冤
k00
k01
k02
k03
k10
k11
k12
k13
···
···
···
···
kN−1,0
kN−1,1
kN−1,2
kN−1,3
冥
is the pressure/frequency coefficient matrix. The row vectors in this matrix show the relationships between pressures and frequencies at various given temperatures, while the column vectors show the relationships between pressures and frequencies at various temperatures. Moreover, these relationships can be expressed in the form of the following cubic fitting polynomial ជ2 + T ជ3共 j = 0, 1, 2, 3兲, . . . . . . . . . . . . (18) kជj = j,0 + j,1Tជ + j,2T j,3 February 2007 SPE Production & Operations
Fig. 8—The effects on coupling distance. 131
System Calibration Experiments. The experiments shown in Fig. 9 were performed to calibrate the telemetry system for accuracy. The system was composed of a wireline-conveyed gauge, a permanently installed gauge, a surface processor and peripheral telemetry components. The accuracies of the various experimental devices are listed in Table 2. The electronic pressure gauges and the components such as the coreless EM coupling and RCequivalent cable were kept in a constant-temperature air bath. The experiments were performed at temperatures of 10, 30, 45, 60, 75, 90, 110, and 125°C. The system was run for at least 2 hours at each fixed temperature to reach thermal equilibrium. The piston-type pressure gauge applied the following approximate pressures to the electronic gauges at each temperature: 0.092, 6.99, 13.88, 20.77, 27.66, 34.56, 41.45, 48.34, 55.23, 62.12, and 70 MPa. The P/T raw data obtained by sensors were converted to the frequencies fជT and fជP , then transmitted to the surface processor through inductive ជ (c , c , c , coupling and equivalent cable. The coefficient vectors C 0 1 2 ជ ជ ជ ជ c3) and the matrix K(k0, k1, k2, k3) in Eqs. 11 and 17 were calculated to fit the actual temperatures and pressures in the temperature range of 0 to 125°C and the pressure range of 0 to 70 MPa. The temperature-fitting error curves for the wireline-conveyed gauge and the permanently installed gauge are distributed symmetrically, as shown in Fig. 10. Their respective maximum errors are approximately –0.09°C at 90°C and –0.3°C at 110°C. The fitted pressure values have an approximately linear relationship with frequencies at each fixed temperature. Fig. 11 shows the fitted pressure curves for the permanently installed gauge, which proves the feasibility of the two-norm least-square curve fitting method. However, the frequency outputs at the fixed pressures vary from 20 to 200 Hz because of temperature changes. These errors, which are mainly caused by temperature drifts of sensors and circuits, are more evident at high pressures where temperature has a much greater effect on pressure. Figs. 12 and 13 reveal the actual pressure-fitting errors for the wireline-conveyed gauge and the permanently installed gauge. Their respective maximum errors are –14.62 kPa at 60°C and 34.56 MPa, and –16.64 kPa at 30°C and 62.12 MPa. These errors result mainly from the response characteristics of the pressure sensor at low temperatures. The experimental calibration results illustrate that the P/T accuracies of the system are respectively less than 0.1% FS and ±0.5°C within the calibrated pressure and temperature ranges.
Fig. 9—Schematic diagram of experimental devices.
gauge, with a pressure accuracy of ±0.02% FS and a temperature accuracy of ±1°C, was installed at a depth of 1644.06 m to monitor the formation information below the valve. The wireline-conveyed gauge was lowered into the hole and positioned at a depth of 1621.2 m to measure the data above the valve. The sonde connected with the gauge acquired data from the memory-pressure gauge by inductive coupling. The tester valve was opened to draw down the pressures in the interval, after which the valve was closed to permit fluid pressures to build up while the measurements were made and recorded. The operation successfully monitored 60 hours of historical data. The data obtained by P/T sensors in the wirelinedeployed gauge were converted to frequencies for remote cable transmission. The P/T frequencies were transferred upward in the “S-P-L-T-L” mode to the surface processor, where they were fitted to compute the actual P/Ts using the fitting coefficients previously determined by the system calibration experiments. The computed fitted data in the surface processor could be displayed and downloaded to the computer through an RS232 serial port. Data recorded by the McAllister gauge were read and used to obtain actual P/T solutions after the gauge was retrieved to the surface. Then the computed P/Ts from the two gauges were analyzed and compared, as shown in Figs. 14 and 15. The curves representing the data from the system developed in this study lie below those obtained from the McAllister gauge data because the P/Ts increase with well depth. The maximum deviations between the two sets of P/Ts are approximately 0.23 MPa and 0.8°C. Assuming a downhole pressure gradient of 1.02 MPa/100 m, a temperature gradient of 2°C/100 m, and a vertical distance of 20 m, the measured P/T deviations are respectively 0.026 MPa and 0.4°C if the data are compared at equivalent well depths. Thus, the system can obtain a pressure accuracy of 0.1% FS in the calibrated range of up to 70 MPa. Although the temperature accuracy of the McAllister gauge is inferior to that of the system, the temperature deviation of 0.4°C
Case Histories. The telemetry system has already been applied to onshore well testing in the Bohai, Huabei, and Jidong oilfields in China. One of the field tests involved an application in the Cha 76-29 well with a well depth of 1746.6 m in the North China oil field. The maximum reservoir temperature was 60°C and the maximum reservoir pressure was 16.4 MPa. The drillstem testing string used a McAllister memory pressure gauge as a standard of comparison for the data to be obtained from the system. The McAllister
Fig. 10—Temperature-fitting error curves. 132
February 2007 SPE Production & Operations
Fig. 11—Pressure-fitting curves at different temperatures.
proves that the temperature information from the system with its accuracy of ±0.5°C is reasonably accurate. Therefore, the system developed here can be employed in the ranges of 125°C and 70 MPa with the calibrated accuracies. Furthermore, the formation pressure decreases sharply when the well is opened initially, then increases rapidly and approaches the static pressure when the well is closed. The formation temperature also decreases or increases accordingly, but the variation amplitude is small. The temperature variation also lags behind the pressure variation every time the well flows and builds up. In addition, an excellent match is observed between the data profiles obtained, confirming the high quality of the system and the feasibility of its use in the field. In comparison with other telemetry methods which experience interference from wellbore storage effects, testing time is significantly reduced, generating considerable savings for the operating company because of the reduced evaluation time. Conclusions A coreless EM coupling-based remote telemetry system for surface readout of bottomhole data during drillstem testing has been applied to onshore well testing. It uses dual electronic gauges to record the pressure and temperature histories above and below a tester valve with valve opening and closing. Data modulated on a power signal are transmitted in the “S-P-L-T-L” mode on the 1-wire. A simplified algorithm implemented on the surface processor is proposed to compute the actual pressures and temperatures for formation transient analysis by fitting the frequency data obtained. A more efficient algorithm can be developed to reduce
Fig. 13—Pressure-fitting error curves of permanently installed gauge. February 2007 SPE Production & Operations
Fig. 12—Pressure-fitting error curves of wireline-conveyed gauge.
the nonlinear errors of pressure sensors and improve data fitting. The experimental results show that coreless inductive coupling can work reliably in spite of the shielding effects of the cylindrical tube. Data transmission has been confirmed at a depth of 1746.6 m with a data rate of one pressure value and one temperature value per second. The system has a calibrated pressure accuracy of 0.1% FS and a temperature accuracy of ±0.5°C and can give the operating company the advantages of real-time availability of bottomhole data at temperatures up to 125°C and pressures up to 70 MPa. Nomenclature ⳱ vector A ⳱ cross-sectional area, L2, m2 bជ ⳱ vector ci ⳱ temperature/frequency coefficient ជ ⳱ temperature/frequency coefficient vector C d1 ⳱ diameter of the small coil L, m d2 ⳱ diameter of the large coil L, m dt ⳱ increment of time t, s dl ⳱ increment of length L, m dB ⳱ increment of magnetic flux mL2t−2q−1, Wb e ⳱ fitting error eជ ⳱ fitting-error vector E ⳱ electric field strength mLt−3q−1, V/m f ⳱ frequency, t−1, s−1 fpi ⳱ pressure frequency,t−1, s−1 fTi ⳱ temperature frequency,t−1, s−1 fx ⳱ frequency of input signal, t−1, s−1 f0 ⳱ frequency of clock signal, t−1, s−1 fជT ⳱ temperature frequency vector ជ fP ⳱ pressure frequency vector
aជ
Fig. 14—Actual pressure curves. 133
References
Fig. 15—Actual temperature curves.
FP ⳱ inverse transfer function of pressure sensor FT ⳱ inverse transfer function of temperature sensor G ⳱ coefficient matrix for the solution of the ជ temperature-fitting coefficient vector C I ⳱ electric current q, A kij ⳱ pressure/frequency coefficient at a fixed temperature Ti kj ⳱ pressure/frequency coefficient at a fixed temperature T kជ ⳱ pressure/frequency coefficient vector at a fixed K l l1 l2 L2 m M n Nx N0 N1 N2 p* i
⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱ ⳱
Pជ ⳱ ជ Pi ⳱ ជ⳱ P* ri ⳱ R⳱ RT ⳱ R0 ⳱ Ti ⳱ T* i ⳱ Tជ ⳱ ជ T* ⳱ Y⳱ ␣1 ⳱ ␣2 ⳱ ជ  j ⳱
temperature T pressure/frequency coefficient matrix length of the coil L, m length of the small coil L, m length of the large coil L, m self-inductance of the large coil, mL2t−2q−2, mH magnetic moment qL2, Am2 coupling factor number of coil turns pulse count of input signal pulse count of clock signal number of turns in the small coil number of turns in the large coil actual fitted-pressure value at the actual fitted temperature value T i*, m/Lt2, Pa fitted-pressure vector at a fixed temperature T fitted-pressure vector at a fixed temperature Ti actual fitted-pressure vector resistance/temperature coefficient resistance, mL2t−3q2, ⍀ resistance at temperature T°C, mL2t−3q2, ⍀ resistance at temperature 0°C, mL2t−3q2, ⍀ temperature, T, °C actual fitted temperature value, T, °C temperature vector actual fitted-temperature vector two-norm least-squares goodness-of-fit expression inductive self-inductance coefficient of the small coil inductive self-inductance coefficient of the large coil vector function fជjT, in which the components are the jth power of the components of vector ជ f T
ជj ⳱ temperature fitting coefficient vector for pressures at different temperatures Subscripts T ⳱ transpose 0, 1, 2 ⳱ the 0th, second, and third power * ⳱ fitted values of actual pressures or temperatures Superscripts p ⳱ pressure T ⳱ temperature 134
Ding, T.H. et al. 2000. Downhole Coreless Electromagnetic Coupling Communication Device. China: Patent No. 00,100,553. Ding, T.H. and Li, C. 2004. Electromagnetic Coupling-Based Downhole Remote Telemetry in Well Testing. Paper presented at the 3rd International Symposium on Instrumentation Science and Technology, Xi’an, China, 18–22 August. Electromagnetic MWD Telemetry System Sets Depth Record Offshore. 2001. Oil & Gas J. 100 (36): 46–47. Harper, G., Almaza, E., Fossa, A., Finley, D., and Strang, G. 2003. Implementation of Advanced Acoustic Telemetry System Adds Value and Efficiency to Well Testing Operations. Paper SPE 80554 presented at the SPE Asia Pacific Oil and Gas Conference, Jakarta, 9–11 September. DOI: http://www.spe.org/elibrary/servlet/spepreview?id⳱80554-MS. Qureshi, Y., Gunarathne, G.P.P., and Christidis, K. 1999. Dynamic Impedance Matching of Transmission Cables for Downhole Tools. IEE Colloquium (Digest) 143: 47–51. Soulier, L. 2003. Method and System for The Transmission of Information by Electromagnetic Waves. U.S. Patent No. 6,628,206. Tochikawa, T., Sakai, T., Taniguchi, R., and Shimoda, T. 1996. Acoustic Telemetry: The New MWD System. Paper SPE 36433 presented at the SPE Annual Technical Conference and Exhibition, Denver, 6–9 October. DOI: http://www.spe.org/elibrary/servlet/spepreview?id⳱ 36433-MS. Trofimenkoff, F.N., Segal, M., Klassen, A., Haslett. J.W., Smallwood, R.E., and Lehner, D. 2000. Characterization of EM Downhole-toSurface Communication Links. IEEE Trans. Geosci. Remote Sensing 38 (6): 2539–2547. Van Riet, E.J., Reitsma, D., and Vandecraen, B. 2004. A Fully Automated System Accurately Controls Downhole Pressure During Drilling. JPT 56 (2): 42–44. SPE-85310-PA. DOI: http://www.spe.org/elibrary/ servlet/spepreview?id⳱85310-PA. Veneruso, A.F. 1990. Apparatus for Electromagnetically Coupling Power and Data Signals Between a First Unit and a Second Unit and in Particular Between Well Bore Apparatus and the Surface. U.S. Patent No. 4,901,069. Veneruso, A.F., Hiron, S., Bhavsar, R., and Bernard, L. 2000. Reliability Qualification Testing for Permanently Installed Wellbore Equipment. Paper SPE 62955 presented at the SPE Annual Technical Conference and Exhibition, Dallas, 1–4 October. DOI: http://www.spe.org/elibrary/ servlet/spepreview?id⳱62955-MS. Whittle, T.M., Lee, J., and Gringarten A.C. 2003. Will Wireline Formation Tests Replace Well Tests? Paper SPE 84086 presented at the SPE Annual Technical Conference and Exhibition, Denver, 5–8 October. DOI: http://www.spe.org/elibrary/servlet/spepreview?id⳱84086-MS.
SI Metric Conversion Factors cycles/sec × 1.0* E+00 ft × 3.048* E–01 °F (°F–32)/1.8 in. × 2.54* E+01 psi × 6.894 757 E+00 psi × 6.894 757 E–03
⳱ ⳱ ⳱ ⳱ ⳱ ⳱
Hz m °C mm kPa MPa
*Conversion factor is exact.
Tianhuai Ding is a professor and thesis director at Tsinghua U., China. His research interests include sensors and intelligent instruments, oil and gas well testing, and optical electronic detection. He plays a key role in designing and developing drillstem testing tools and is in charge of project and technology integration for telemetry systems. From 1984 to 1986, he worked as an associate researcher at the U. of Munich, Germany. He holds a BS degree in precision instrumentation from Tsinghua U. Li Cheng is currently a PhD candidate in science and technology instrumentation at Tsinghua U. His research interests include electronic sensors and industrial automation and instrumentation systems related to oil and gas well testing. Cheng holds BS and MS degrees in electromechanical engineering from Hebei U. of Technology, China. February 2007 SPE Production & Operations
