Tube Expansion Research Papers

Plugging criteria for thinned high-pressure feedwater heater tubes considering pressure and thermal loading

2024, Journal of Mechanical Science and Technology

When heat exchanger tubes have critical thinning which could cause unsafe conditions, it is general practice to plug the tubes so that they are out of service to prevent rupture failure at the thinnest location. Plugging criteria for the... more

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Lean-and-Green Strength Performance Optimization of a Tube-to-Tubesheet Joint for a Shell-and-Tube Heat Exchanger Using Taguchi Methods and Random Forests

by Panagiotis Boulougouras

2024, Processes

This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY

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Lean-and-Green Strength Performance Optimization of a Tube-to-Tubesheet Joint for a Shell-and-Tube Heat Exchanger Using Taguchi Methods and Random Forests

by Panagiotis Boulougouras

2024, Processes

The failing tube-to-tubesheet joint is identified as a primary quality defect in the fabrication of a shell-and-tube heat exchanger. Operating in conditions of high pressure and temperature, a shell-and-tube heat exchanger may be... more

Table 3. Chemical composition % and mechanical properties of steel P265GH.

Next, using a CAD/CAM mechanical design software (Fusion 360, Autodesk, San Francisco, CA, USA), 3D solid models were designed in order to generate the required G-code programs for the CNC milling treatment of the test tubesheet specimens (Figure 3). For the machining of the test tubesheet specimens, a 3-axis vertical milling CNC center (Genos M560-V-e, Okuma, Oguchi, Aichi, Japan) was employed which was also integrated with an OSP-P300MA-e (Okuma, Oguchi, Aichi, Japan) controller as depicted in Figure 5. Table 4. Chemical composition % and mechanical properties of steel P215NL.

Figure 5. The CNC machining center with the integrated controller. Figure 4. (a) Drill inserts and drill-body tool-holder; (b) drill inserts, boring tool-holder, and finish boring head adapter; (c) T-slot milling cutters; (d) tubesheet specimens after exiting the machin- ing process.

Figure 6. (a) Tube expanding machine; (b) three-roller tube expander.

Because of the criticality of a proper specimens’ preparation in ensuring trustworthy test-joint strength measurements, the grooves on the tubesheet-hole walls, along with the tube inner and outer surfaces, were pre-treated to rid of any surface blemishes, such as burrs, grease, oil residue, rust, or dirt. The finalized torque settings which were applied during the tube-rolling process and their expected tube wall thickness reduction percentage are listed in Table 6. Before conducting the tube expansion process (Figure 7d), the required tube wal reduction percentage for each experiment was pre-evaluated to select suitable trial setting: for the OA scheme. The relative tabulations from refs. [51,52] were adopted to complet the ensuing detailed calculations. Equation (1) from ref. [7] was used to allow a realistic pre-estimation of the tube inner diameters for the tube-rolling process (Table 5). Moreove! the tubesheet hole diameter, as well as the tube inner and outer diameters, were measurec thrice each (Figure 7), for every replicated nine-run set of experiments, and their mear values were recorded. The tolerance limits of individual experiment were determined to be the expanded tube inner diameters for eact within +£0.05 mm.

Before commencing the tube push-out testing procedure, the tube ends of the tube spec- imens were sealed with welded plugs as depicted in Figure 8a. A tensile/compressive test- ing machine (CTM/50, Galdabini, Cardano al Campo, Varese, Italy) was utilized (Figure 8b) to perform the tube push-out tests. The testing machine was integrated with GraphWork 5.0 test software which recorded the applied loads on the tubes during the push-out testing process. Moreover, a round steel bar of material S355 (Figure 8c), with nominal diameter and length dimensions of 14 mm and 295 mm, respectively, was inserted into the tube end to direct the axial load transmission (Figure 8d); it forced the tube out of the tubesheet hole to the point of mechanical failure of the joint or the tube (Figure 8e,f). The push-out tests were carried out at a constant crosshead speed of 5 mm/min. After the axial tube- displacement process was initiated, the applied load was monitored and recorded by the esting equipment. The maximum load was applied at the start of the displacement cycle and then it was gradually decreased. The terminal condition of the tube-to-tubesheet speci- mens exiting the push-out testing is demonstrated in Figure 8g. Duplicate joint strength measurements were obtained for the nine-run tube-to-tubesheet specimens.

Figure 8. (a) Tube end sealing; (b) push-out testing machine; (c) round S355 steel bar; (d) preparation for push-out testing; (e) before tube displacement; (f) after tube displacement; (g) tested tube-to- tubesheet specimens after tube push-out testing completed.

Figure 9. The input-output saturated L9(3*) OA arrangement for duplicated response data.

Figure 10. The two separate data reduction steps to generate the mean (M) and SNR responses, respectively.

Figure 11. (a) Response graph for joint strength SNR; (b) response graph for joint strength mean

The optimal layout of the specimens (the 10th factorial recipe), which is prepared on the suggested parameter adjustments from the previous subsection, is re-drawn and is shown in Figure 12. The confirmation run was comprised of five pairs of machined tubes and tubesheets, which are shown in Figure 13. Thus, the confirmation run was replicated five times. The collected measurements from the confirmation experiment are listed in Table 16. Figure 12. Drawing of the optimal tube-to-tubesheet specimens to be used for the confirmation runs.

Figure 13. Tubes and tubesheet specimens machined for the confirmation run.

Figure 14. Cont. A post-screening assessment could be initiated by depicting the four controlling factors in boxplots for each individual controlling factor to emphasize the visual aspect of their variability owing to different settings. In Figure 14, the corresponding four boxplots have been prepared (MINITAB 19.0, State College, PA, USA). It is immediately discerned that only the tube wall thickness reduction could amass distinguishable variation across settings. A linear dependence emerges which is solely framed by the two operating endpoints of this factor. The second (middle) setting merely overlaps across both end settings. The rest of the controlling factors do not pass the test of visual evidence. Perhaps, their influence on the joint strength response is dubious; their confidence intervals for their medians are overlapping across all settings. Equally striking is the asymmetric location tendencies in the data sample median values as exposed by individual settings.

Figure 14. Boxplots for the joint strength (JS in N) due to: (A) clearance (A in mm), (B) number of grooves (B), (C) groove depth (C in mm), and (D) tube wall thickness reduction (D in %).

No particular side is favored in the numerical accumulation of the observations; t hey are predominantly asymmetric across all four factors. A plausible question would be whether any additional trial runs would refine more the precision around the magnitude estimation of the effects and, hence, would assist to expediting the practical aspect of the profiling process. To check consistency in the duplicated trials of the joint strength experiments, the two replicates are regressed against each other (Figure 15). It results in a symmetrical line fitting—the fitted line slope is 1.03—and the adjusted coefficient of determination is estimated at 98.7%. It is inferred that there is no immediate concern to conduct more replications or selectively augment the volume of observations. It is useful to also obtain an alternative prediction from a classic algorithmic solver such as the Random Forest. The trial outcomes from running the Random Forest module 10 times are summarized and tabulated in Table 17 (IBM SPSS v.29, Armonk, NY, USA). The precision of the predictions is deemed adequate. This was determined by evaluating the descriptive statistics for a start-up sample. The Random Forest quantities of statistical interest were the mean of squared residuals (MSR), the percentage of the explained variation (VE), and the mean squared error (MSE) for the four controlling factors. Then, the descriptive statistics were re-evaluated by adding in the final sample (n = 10), an additional batch of five prediction runs from executing the Random Forest module. The mean difference estimates between the two sampling phases were found to converge to the listed precision with an experimental test power of 0.95. Thus, no further Random Forest solver runs were deemed necessary.

Figure 16. Error convergence according to the tree number selection in the Random Forest routine execution. MSR = Mean of Squared Residuals, VE = Variation Explained (%), MSE = Mean Squared Error: A = Clearance, B = Number of Grooves, C = Groove Depth, D = Tube Wall Thickness Reduction.

Figure 17. Minimization of the out-of-bag error (OOB) due to number of variables randomly samplec as candidates at each split (mtry) in the Random Forest solver. sampling tactic, keeping replication to a minimum and finally performing confirmation runs to a restricted number. Nevertheless, a conceived factor such as the surface finish of the tubesheet holes may induce a detectable effect on joint strength [7,9]. This factor was not examined in the present research. According to the TEMA standards [7], it is stated: “a smooth tubesheet hole can provide a pressure tight joint at a lower level of wall reduction, but a rough tubesheet hole provides more mechanical strength than a smooth tubesheet hole” (pp. 10-31). Since rough tubesheet holes and strong joints may be related, extending his research would uncover valuable links to enhancing the joint integrity on the particular production materials as used in this study.

Figure 18. Factorial profiling using the Random Forest solver to evaluate the partial dependence of: (1) Clearance (A), (2) Number of Grooves (B), (3) Groove Depth (C), and (4) Tube Wall Thickness Reduction (D).

Table 1. Control factors and their associated settings. As stated earlier, fast recipe-making is imperative when concocting industrial trials. The standard Taguchi-type L9(3*) OA [43] accommodates as many as four controlling factors by requiring nine specific experimental recipes to be executed. The four factors are accommodated in a saturated scheme. Saturated experimental plans may also be deemed ‘lean-and-green’, because they promote the maximization of the investigated influences. Each factor participates in the experimental plan at three operating settings through the preset configuration of the OA sampler. The promoted ‘lean-and-green’ engineering aspect is evident at this point, since the total work volume is reduced by a ratio of 1/9; a full- factorial arrangement of the four 3-level factors would otherwise demand a total of 81 (=34) trials. In Table 2, we list the scheduled factorial combinations according to the Lo(3*) OA arrangement. The nine-run OA set-up was makeshift replicated twice. This is another minimal (lean-and-green) replication practice. Finally, the tolerance limits for both the clearance and the groove depth were decided to be restricted within +0.02 mm.

Table 2. Four-factor experimental arrangement Lo (34) OA for testing joint strength.

Table 6. Torque values used for tube rolling. 2.5. Joint Strength Measurements

Table 7. Duplicated dataset, mean, and SNR estimations of the joint strength.

Table 8. Response table for joint strength mean (N). The remaining two factors, the clearance and the groove depth, are behaving as of a lesser importance—at a similar effect magnitude between them. According to the response graph for the joint strength mean, the optimal setting of the tube wall thickness reduction is 12% and the respective optimal setting of the number of grooves is three. Similar observations may be drawn from the response table (Table 9) for the joint strength SNR. However, the discrepancies between settings are relatively smaller for the SNR estimations. The tube wall thickness reduction and the number of the grooves elicit a maximum difference in the joint strength mean SNR values of 6.53 dB and 4.55 dB, respectively. Overall, it may be remarked that both factors cause minor disturbance on the mean SNR response of the joint strength. Similarly, the main effects plot for the joint strength mean SNR shows that the corresponding optimal settings are exhibiting the same pattern as in the main effects for the means plot, i.e., optimally set at 12% for the former and three for the latter.

Table 9. Response table for the joint strength SNR (dB). Tt o proceed with the analysis part, the statistical significance of the outcomes is obtained the ANOVA treatment. As expected, the two data reduction steps have each separately sized the duplicated form of the original datasets—mean and SNR estimations—to a ‘unreplicated response’ form. Additionally, the experimental OA-plan was decided fully loaded with controlling factors (‘saturated’ condition). Hence, the usefulness of the ANOVA is restricted to offering only a relative strength evaluation among the factoria butions. This is because in the saturated-unreplicated form, an OA sampler does not permit significance estimations in the absence of a residual error. This is observed for means and SNR estimations (Tables 10 and 11). Despite that, till worthwhile to contemplate the leading contributions, in terms of mean sum of from down virtua to be contri in bot. itis s h ANOVA tables, squares calculations. Ac sum of the mean sums of squares for both estimator treatments—mean and SNR joint strength response. Simi sums of squares of the Even though there is no the goodness-of-fit computations to assess the status of the predictions, nevertheless, a qualitative comparison between the data-reduced (mean and SNR) joint strength variable with the simplistic response graph and response table results seems to lead to an agreement. tually, the tube wall thickness contributes about 55% to the tota arly, the number of grooves contributes to the mean and SNR joint strength response about 30.57% and 25.74%, respectively. available information from either the statistical significance or

Table 10. ANOVA treatment for the response means of the joint strength.

Table 11. ANOVA treatment for the response SNR estimations of the joint strength. *: no value available.

A = Clearance, B = Number of Grooves, C = Groove Depth, D = Tube Wall Thickness Reduction. Table 12. Stepwise ANOVA results for unprocessed response joint strength dataset.

A = Clearance, B = Number of Grooves, C = Groove Depth, D = Tube Wall Thickness Reduction. Table 13. Stepwise regression coefficients for the joint strength response dataset.

A = Clearance, B = Number of Grooves, C = Groove Depth, D = Tube Wall Thickness Reduction. Table 14. Best subsets regression (B and D participate in all models).

Table 15. Predicted values for the maximum performance of the joint strength. 3.3. Confirmation Run for the Optimal Parameter Settings

Table 16. Joint strength observations from the confirmation trial runs. Even though only five confirmation runs were programmed to constitute an initial test sample, the tapered 95%-confidence-interval estimation (Table 16) hinted that no additional confirmation-run samples would be meaningful at this stage. The notion of lean thinking also extends to minimizing the confirmation trials, if the trends are already practically explainable and there is a firm grasp on how joint strength variability is distributed among its accounted sources. Moreover, this was attested because of the accompanying very low value of the coefficient of variation of the confirmed joint strength performance (0.0098). However, the discrepancy between the predictions due to the Taguchi method and the stepwise regression analysis, on one hand, and the confirmation-run joint strength mean response values, on the other, was calculated at a relative percentage of about 14%. While he complexity of the physical phenomena behind this mechanical process is known to be high, and in part, it could justify such realistic deviation, particularly under the espoused condition of the ‘lean-green-driven’ (miniaturized) experimentation, in the Discussion section there will be further exploration about the cause of such discrepancies. Conversely, the confirmation-run joint strength performance matches fairly close to the maximum observable joint strength measurement in the OA trials which was found to be 45,371.25 N Table 7). Thus, their difference may be viewed now more tolerable at the relative percentage of 3.3%. It is concluded that this study might have produced useful empirical relationships, but an alternative non-statistical treatment might provide insight from a different angle. This last issue will be discussed in the next section.

Table 17. Random Forest trial statistics to fit the saturated Lo(3*) OA joint strength dataset.

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Impulse facility simulation of hypervelocity radiating flows

by michael macrossan

2023, Science & Engineering Faculty

We describe the X-series impulse facilities at The University of Queensland and show that they can produce useful high speed flows of relevance to the study of high temperature radiating flow flields characteristic of atmospheric entry.... more

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Coupling and Damping Effects on the Dynamics of Submerged Expanded Tubes in Borehole Wells

by Ali Karrech

2023, The Journal of Engineering Research [TJER]

Hydraulic expansion of submerged tubes is accomplished by propelling a mandrel through it using differential pressure. This process deforms the tube beyond its elastic limit. Toward the end of the expansion process, the mandrel pops out... more

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Plugging criteria for thinned high-pressure feedwater heater tubes considering pressure and thermal loading

by Hồng Vấn Đinh

2023, Journal of Mechanical Science and Technology

When heat exchanger tubes have critical thinning which could cause unsafe conditions, it is general practice to plug the tubes so that they are out of service to prevent rupture failure at the thinnest location. Plugging criteria for the... more

Fig. 1. Modeling of straight tube section of de-superheating zone con- nected to the U-tube section (R is boundary condition from U-tube).

Fig. 2. Types of thinning for feedwater heater tubes. 2.3.1.2 Thermal loading types: Uniform, eccentric, and local thinning. Fig. 2 gives further details about the types of thinning in the tubes.

2.3.2 Eccentric thinning Consequently, three principal stresses for the case of uni- form thinning are shown in Eqs. (11)-(13). Using a bipolar coordinate system, an analytical solution was derived for the stresses of an internally pressurized eccen- tric pipe [11]. For an eccentric tube subjected to internal pres- sure, the maximum hoop stress is generated at the inner wall at the thinnest location. The maximum hoop stress, which is the first principal stress, was derived as Eq. (14). The allowable stress to prevent failure was defined by the proof stress which is the mean of the yield strength and ulti- mate tensile strength of the material.

Fig. 3. Stress distribution for uniform and eccentric thinning of the tube determined by FEA under combination of thermal and pressure loading

Fig. 5. Comparison of the stresses between uniform thinning and ec- centric thinning: (a) Under internal pressure loading only; (b) under a combination of pressure loading and thermal loading. though the thinning modes are different.

Fig. 6. (a) Stress distribution under combined thermal and pressure loading with respect to radial coordinate; (b) stress at inner surface as a function of thinning rate.

Table 1. Operating conditions of three feedwater heaters.

Table 2. Material properties of SA-213 TP304N.

Table 3. Tresca stress results for various thinning conditions for the cases of uniform and eccentric thinning.

Table 4. Plugging criteria for outer thinning using Tresca and Von Mises criteria. the difference in stress results between the uniform and eccen- tric thinning models was insignificant (less than 10 %). The two models have practically same stress when a thickness of uniform thinning equal to the smallest thickness of the eccen- tric thinning is employed in calculation. Therefore, we can use the equation for the uniform thinning, Eq. (11) (with the minimum thickness value in the eccentric thinning), instead of the equation for the eccentric thinning model, Eq. (14), for determining the minimum required thickness for plugging criteria.

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Near-resonant holographic interferometry of hypersonic flow

by Halina Rubinsztein-Dunlop

2023, Shock Waves

Near-Resonant Holographic Interferometry is a powerful technique which extends the established advantages of conventional holographic interferometry by allowing a species-specific number density to be determined. It has been tested in the... more

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Flow tagging velocimetry in a superorbital expansion tube

by Halina Rubinsztein-Dunlop

2023, Shock Waves

A non-intrusive laser-based method for direct velocity measurements has been demonstrated in a superorbital flow facility. The method is based upon laser enhanced ionisation velocimetry in which a tagged region is created by two step... more

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Trends in the Development of Machinery and Associated Technology

by Saif Al-Hiddabi

2022

A model to study the dynamics of a stick-slip phenomenon in expansion of tubes was developed. During the permanent deformation process, the system experiences large friction forces at mandrel/tubular interface. This result in an increase... more

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Indirect Extrusion: A Multifaceted Approach of Sub-surface Tubular Expansion

by Zahid Qamar, Sayyad

2022, Extrusion of Metals, Polymers and Food Products

Extrusion and indirect extrusion is a very old manufacturing process used in multitudes of applications mainly focused on transportation, household and power industries. Indirect extrusion has found an interesting application in petroleum... more

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On the performance analysis of AHSS with an application to SET technology – FEM simulations and experimental measurements

by Zahid Qamar, Sayyad

2022, Thin-Walled Structures

Solid expandable tubular (SET) is an innovative breakthrough in the petroleum industry that aims to resolve issues associated with deep wells. It consists of in-place plastic deformation of tubular diameter resulting in larger conduit... more

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Stick− Slip Friction Modeling in Tube Expansion

by Zahid Qamar, Sayyad

2022

A model to study the dynamics of a stick-slip phenomenon in expansion of tubes was developed. During the permanent deformation process, the system experiences large friction forces at mandrel/tubular interface. This result in an increase... more

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Tube Expansion Under Various Down-Hole End Conditions

by Omar Al-Abri

2022, The Journal of Engineering Research [TJER]

Fossil hydrocarbons are indispensables commodities that motorize the global economy, and oil and gas are two of those conventional fuels that have been extracted and processed for over a century. During last decade, operators face... more

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Effects of expansion rate on plasticity and structural integrity of downhole tubular

by Nashmi Alrasheedi

2022, International Journal of Pressure Vessels and Piping

The current work presents an investigation of the effects of expansion rate on plasticity and structural integrity of down-hole solid tubular. Down-hole tubular has proven itself to be a promising technology in oil well applications by... more

Table 1: Geometric parameters of mandrel and tubular for 16%, 20%, and 24% ER Figure 3: Stress strain behavior of tubular material subjected to uniaxial tensile load

Table 1: Contact pressure at TMI for different mandrel velocities

premature failure of the tubular, since the contact pressure at both regions is below ultimate strength. These are useful results for the operators and engineers who desire to expand the tubular with higher mandrel velocities to reduce well completion time. Since the contact pressure at tubular-mandrel interface is found to be high in most of the cases, it is most likely that the expanded tubular may have residual stresses. Figure 7 shows the variation of equivalent stress with respect to mandrel position for 16%, 20%, and 24% expansion ratio (ER), while Table 1 shows the figurative values of the same.

Figure 9 and Table 4 show the variation of expansion force at different mandrel velocities for 16%, 20%, and 24% ER. A prominent variation in expansion force is observed for different expansion ratios, while a minor variation is found at different velocities with similar expansion ratio. Maximum magnitude of expansion force is found at 15 mm/sec for all expansion ratios 1.e. 1.00839, 1.09150, and 1.19828 MN at 16%, 20%, and 24%, respectively. The results conclude that velocity of mandrel has minor or negligible effects on expansion force.

Table 4: Expansion force variation with mandrel position at different mandrel velocities

Figure 10: Variation of thickness with respect to tubular length at 16%, 20%, and 24% ER

Figure 11 shows shortening of tubular length against mandrel position for different mandrel velocities and expansion ratios. For simplicity, results are only shown for 5 mm/sec mandrel velocities at 16%, 20%, and 24% expansion ratio. It is noted from Table 7 that as for thickness reduction, there is negligible effect of mandrel velocity on length shortening. Nevertheless, length shortening is a critical post-expansion property of tubular, which becomes very important when two consecutive expanded tubulars are assembled in a well. Special attention must be paid in estimating this characteristic of tubular after an expansion process.

It is clear from Figure 6(b) that a significant amount of contact pressure develops between tubular and mandrel at two locations, i.e. close to the beginning, region 1, and at the end, region 2, of TMI. It means that the interaction between tubular and mandrel is a two-point contact and may cause bending of inclined surface, resulting in outward deflection in radial direction of tubular. This phenomenon is also discussed in [22]. The two pressure peaks at the interface may result in stick-slip phenomenon and might reduce burst and collapse strengths of tubular due to excessive thinning of wall thickness. It is also evident from simulation results that the velocity of mandrel does affect the contact pressure but no definite trend found in the results. The maximum contact pressure of 777 MPa is noticed at 15 mm/sec for 16% ER, while the lowest is 347 MPa at the same velocity. It is also evident from Table | that contact pressure (at region 2) at 15, and 20 mm/sec is almost equal to the average yield strength of tubular material (oy = 625 MPa), while at region | for 15, 20, and 25 mm/sec, it is equal or beyond the average ultimate strength of tubular material (dy7; = 710 MPa), respectively. This means that the propagation of mandrel at velocities 15, 20, and 25 mm/sec may cause failure due to excessive thinning of the tubular wall thickness. Furthermore, expanding the tubular at 5 and 10 mm/sec would not cause any

Table 2: Equivalent stress at inner surface of tubular at different mandrel velocities It is obvious from Figure 7 and Table 2 that expanded tubular in all cases of expansion ratios and mandrel velocities is not stress free. The residual stress varies from 150 to 247 MPa, which is considered as high magnitude of stress. The residue of stress is found to be small at highest velocity and large at lowest. These results are in congruence with material response subjected to different strain rates. The knowledge of residual stress in the tubular is crucial before implementing it into application where the presence of residual stress may reduce burst and collapse strengths of tubular. Figure 8 shows variation in the magnitude of equivalent plastic strain at 5 mm/sec mandrel

Table 3: Equivalent plastic strain at inner surface of tubular at different mandrel velocities A thorough understanding of other post-expansion characteristics of tubular such as expansion force, thickness variation, length reduction, and surplus deformation are also important parameters, which are required in designing of oil and gas wells. These parameters are needed by engineers and operators to exactly estimate the length and thickness clearances for perfectly assembling different tubulars together as well as in determining the structural integrity of tubular.

Table 5: Variation of thickness at different mandrel velocities Figure 10 shows the variation in tubular thickness with respect to mandrel position for 16% 20%, and 24% ER at 5 mm/sec mandrel velocity while Table 5 represents the percentags thickness reduction at different expansion ratios and mandrel velocities. As is obvious from thes results, the reduction in thickness varies slightly with the change in mandrel velocity. Maximun thickness reduction is found at 25 mm/sec for all expansion ratios. This may infer that th velocity of mandrel has induced some effects on the thickness reduction. It could be estimatec that the effect will be more pronounced at velocities higher than 25 mm/sec. However, for th velocity range considered in this work, no significant effect of mandrel velocity is observed o1 thickness variation of tubular. The results of thickness reduction, Table 5, can be utilized t compute burst and collapse strengths of down-hole tubular through API equations [25].

Table 6: Burst and collapse strengths of tubular at different mandrel velocities and expansion ratios Table 6 clearly indicates that the expanding of tubular at different mandrel velocities and expansion ratios may result in higher or lower burst and collapse strengths. Higher value of burst strength is required to avoid failure in the presence of high internal pressure, while a higher value of collapse strength is needed in order to prevent failure from surrounding rock loads, which can result in collapse of an oil well. It is evident from Table 6 that tubular expanded with 16% expansion ratio has highest average burst and collapse strengths. The variation is burst and collapse strengths are noted for 16% and 20% ER at different velocities, but they are almost constant for all velocities at 24% ER. This shows that tubular burst and collapse strengths do not significantly change with respect to mandrel velocity at 24% ER. This is due to the fact that at higher expansion ratio, tubular material is significantly strain hardening, so there is negligible change in thickness of tubular. Therefore, at higher expansion ratio, burst and collapse strengths do not affect by the strain rate. The results reported in Table 6 are very useful for field engineers to estimate post-expansion burst and collapse strengths of tubular under different expansion ratio and strain rates. Moreover, field engineers could make a decision as to which mandrel velocity

Table 7: Percentage length shortening at different mandrel velocities Expansion tests for 174.6 mm inner diameter and 9.5 mm wall thickness tubular have been done for 16%, 20% and 24% expansion ratios. A finite element model for tubular-mandrel system has been developed using commercial software ABAQUS and has been validated through experimental observations. FE model is then used for simulations of tubular expansion to study

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“Design and Analysis of NTIW TUBESHEET for TEMA type shell and tube type heat exchanger” tubesheet compare stress result different TUBESHEET TEMA STD and Analysis and evaluate overstress region in NTIW TUBESHEET)

by MOHAMMAD JIVA

2022

Tubesheet of the Shell and tube heat exchangers designed as per part UHX of ASME Pressure vessels codes. In the past, the tubesheet was designed by this analytical method irrespective of tube layout. However this analytical method does... more

Figure 1 solid model of ASME heat exchanger Figure 2 solid model of TEMA heat exchanger

Following Fig. 3 and 4 show meshed model of ASME and TEMA model of tube sheet and its adjacent components. Sweep meshing has been applied to tube sheets, baffles and tubes. The shell and both channels are meshed by using mapped meshing. As the tubes are welded to the tube sheet so the nodes of the tubes are merged with nodes of the tube sheet surface area. Figure 4 meshed model of TEMA heat exchanger Figure 3 meshed model of ASME heat exchanger

Following Fig. 7 indicates mechanical boundary conditions applied to models for obtaining solution, It is. noted from design procedure that the most critical loading condition on tube sheet is shell side pressure acting’ with thermal expansion. So, heat exchanger model were analysed for this loading condition. Shell side pressure is applied.at the inside of shell, across the baffles, outside surface of tubes, inside surface of tube sheet. The differential thermal expansion occurs between shell and tubes due to temperature difference between them. Since 1/4th portion of model is prepared, symmetrical boundary conditions have been applied at the symmetrical plane as shown below. Following Table V shows the-mechanical loads and displacement applied on heat i a The flue gases are in direct contact with the inlet and outlet channel from inside so there is no need to provide heat transfer coefficient. All the data of heat transfer coefficient are referred from “Heat Transfer and Heat Exchanger” Text Book

Fig.7 mechanical boundary condition applied on heat exchanger

The results obtained from analysis are much less than that obtained from code design. So the tube sheet thickness can be more optimized by FEA methods and FEA results can be used as an optimization tools for design. In both codes there is no provision of tube sheet with knuckle design. The stresses in tube sheet with knuckle are less than the flat circular plate placed on elastic

As per TEMA code the tube s heet thickness at maximum effective pressure is 73.95mm and assumed thickness is 75mm which is within tolerance of + 1.5. From calculation results according to UHX code the most critical load case is load case 6 that is shell side pressure acting along with dif ferential thermal expansion. The net effective pressure acting on tube sheet as per UHX code is 2.28 MPa and 1.073 MPa as per TEMA. The net effective pressure acting on tube sheet in UHX design is more than the TEMA because UHX code provides formula for pressure acting on rim portion of tube sheet separately which was then included for calculating net effective pressure acting on tu the solid portion of the tube s bending of tube sheet. be sheet. This is one of the important factor which effects on bending stress (1) in tube sheet. Rim is heet on which both shell side and tube side pressures are acting and theses pressure contributes on

© 2021 IJCRT | Volume 9, Issue 1 January 2021 | ISSN: 2320-2882

For this special heat exchanger application of ASME Sect. VIII Div. | and TEMA design procedure have been successfully implemented which was then analysed by using FEA software package ANSYS 13. The setting up FE model is time consuming, but the results may be worth the effort because from result Table VI the bending stress in tube sheet by finite element analysis are very less than stress calculated by both codes. The factor of safety used for designing the tube sheet in both standards are more which increases dimensions of the heat exchanger components. It is concluded that the tube sheet thickness of given heat exchanger is safe according to the codes and analysis.

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Structural behavior of solid expandable tubular undergoes radial expansion process – Analytical, numerical, and experimental approaches

by Omar Al-Abri

2022, International Journal of Solids and Structures

Today's structures have to meet increasingly rigorous requirements during operation. The economic and human costs of failure during service impose a great responsibility on organizations and individuals who develop new products as well as... more

Today's structures have to meet increasingly rigorous requirements during operation. The economic and human costs of failure during service impose a great responsibility on organizations and individuals who develop new products as well as those who select/integrate products in a final engineering design. A crucial aspect for successful product development and/or inclusion is the careful selection of the best material(s), derived from an informed awareness of the capabilities and opportunities afforded by all candidate materials, together with a design that takes full benefit of those competencies. Thick-wall tubular is an example where all these issues are playing a major role in deciding their industrial applications. Given for their desirable features of high strength and geometrical shape, they are widely used in aerospace, marine, military, automotive, oil and gas, and many other fields. This paper focuses on developing analytical solution to investigate the structural response of thick-wall tubulars undergo plastic deformation due to expanding them using a rigid mandrel of conical shape. Volume incompressible condition together with the Levy-Mises flow rule were used to develop the equations which relate the expansion ratio of the tubular to the length and thickness variations. Besides, Tresca's yield criterion was used to include the plastic behavior of the tubular material. Further to this, a numerical model of the tubular expansion process was also developed using the commercial finite element software ABAQUS. Experiments of tubular expansion have been conducted using a full-scale test-rig in the Engineering Research Laboratory at Sultan Qaboos University to validate the analytical and numerical solutions. The developed analytical and numerical models are capable of predicting the stress field in the expansion zone, the force required for expansion, as well as the length and thickness variations induced in the tubular due to the expansion process. Comparison between analytical, experimental, and simulation results showed that a good agreement has been attained for various parameters.

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Hybrid sliding PID controller for torsional vibrations mitigation in rotary drilling systems

by fi3a fi3och

2022, Indonesian Journal of Electrical Engineering and Computer Science

During the drilling process, the drilling system devices can be exposed to several types of phenomena incited by lateral, axial, and torsional vibrations. The latter can lead to severe damages if they are not efficiently controlled and... more

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Reliability Analysis of Vertical Well Casing: Comparison of API 5C3 and ISO 10400

by Amirah Husna Abdul Halim

2022

1Civil & Environmental Engineering Department, Universiti Teknologi PETRONAS, 32610 Perak, Malaysia Abstract. Well casings are designed against burst strength pressure. The same formula is used in determining the remaining burst strength... more

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Stress Reduction in Thick-walled Cylinders under Internal Pressure

by M.R. Hematiyan

2022

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Numerical simulations of casing collapse: a case study in the South of Iran

by Raoof Gholami

2022, Petroleum and Mineral Resources

Collapse of casings during production has been reported on many occasions in different fields. Regardless of the root for the failure of the casings, this is due to stresses applied to the casing exceeding its ultimate strength. Failure... more

The field is producing for about 60 years now. The first casing collapse in this field was reported in one well in 1974. Since then, collapse of 48 wells drilled in he field has been reported indicating the need for further investigation of this problem. Figure | shows plot of number of wells experienced casing collapse each year. This figure shows that there is no trend indicating that the collapse has been due to the age of the pipe or bad cementing considering that similar casing pipes has been used and cementing jobs were done in a similar way in all wells. Past studies by many researchers suggest that there is no single mechanism responsible for the collapse of casing in this field but it due to combination of different mechanisms [2].

Model geometry before (left) and after (right) applying external uniform pressure.

Displacement of casing wall at an external pressure of SOMPa.

Von-Mises stress distribution of casing wall at an external pressure of SOMPa.

Comparison of collapse pressure in casings with different D/t ratios.

therefore different casing collapse data is presented with a variation of yield strength depending on specific application. In presented Tables the ratio of diameter to thickness of test specimens was between 10 and 40. To measure the steel imperfection, the specimens were sealed at both ends and placed in a high- pressure test facility. The collapse experiments were carried out under volume controlled loading conditions. The applied external pressure was monitored using a calibrated pressure transducer and analogue time-based recording equipment. The pressurization rate was low enough to avoid strain rate effects. The geometry and stiffness of casings used for experimental study were carefully imported into the ABAQUS numerical modeling program. Figure 6 compares the results of FEM and those of the experiments. Plastic Strain (PE) values computed according to the external pressure applied on the outer sides of the casing.

Comparison of numerical simulation results with experimental data.

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Parametric Analysis of the Mandrel Geometrical Data in a Cold Expansion Process of Small Holes Drilled in Thick Plates

by Tiziana Segreto

2022, Materials

Cold expansion technology is a cold-forming process widely used in aeronautics to extend the fatigue life of riveted and bolted holes. During this process, an oversized mandrel is pushed through the hole in order to yield it and generate... more

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Stick-slip analysis in vibrating two-layer beams with frictional interface

by Kourosh heidari shirazi

2022, Latin American Journal of Solids and Structures

This paper attempts to present a new analysis for dynamical behavior of two-layer beams with frictional interface which held together in a pressurized environment, including stick-slip nonlinear phenomenon. To achieve a proper outlook of... more

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Stress Concentration Factor of Expanded Aluminum Tubes Using Finite Element Modeling

by Dr. Abdennour Seibi

2022

This technology subjects solid tubes and sand control screens to large radial plastic deformations through expansion. The expansion process is usually performed either mechanically or hydraulically using conical mandrels. Tubular (casing)... more

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Effect of Cold Work on Mechanical Properties of Expandable Tubulars

by Zahid Qamar, Sayyad

2022

Some of the challenges being faced by the petroleum development industry worldwide are low-cost and more efficient well completion, water-shutoff, other types of zonal isolation, etc. Successfully addressing these issues, solid expandable... more

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Stick−slip Friction Modeling in Tube Expansion

by Omar Al-Abri

2022

A model to study the dynamics of a stick-slip phenomenon in expansion of tubes was developed. During the permanent deformation process, the system experiences large friction forces at mandrel/tubular interface. This result in an increase... more

Figure 1. Friction Models: (a) Classical Model, and (b) Karnopp’s Model Expansion process induces wall thickness imperfections due to local plastic deformation as a result of stick and slip of the mandrel. When the mandrel starts moving through the tubular, the chance of sticking to the expanded tubular increases due to lack of proper lubrication, tubular surface irregularities, and the presence of welded and/or threaded connections. This means that higher force is required to overcome the excess of friction and push the mandrel forward. When the applied force reaches the limit to overcome maximum static friction, the mandrel tends to slip relative to the expanded tubular. This phenomenon is known as a “stick-slip” phenomenon. The study of stick-slip vibrations is faced with difficulties, as during the stick-slip motion two different mechanisms are taking place; the motion is governed by a static friction force in the stick phase and a velocity dependent kinetic friction force in the slip phase. Modelling of these two mechanisms yields a set of differential equations with discontinuous region as shown in Figure 1(a). A standard method to solve this discontinuity problem is by applying a smoothing method. The classical switch model using a switching method between slip, stick and transition phases poses the discontinuity problem. Hence Karnopp [6] model that is using a zero velocity interval to define slip, stick and transition phases is used as shown in Figure 1(b). 3. GOVERNING EQUATION Assuming thin wall tubular, conical shape mandrel to expand the tubular, constant rate of deformation and uniform contact pressure at tubular/mandrel interface, the governing equations can be derived using equilibrium and incompressibility conditions. An infinitesimal element of the expanded tubular wall subjected to contact pressure and friction as well as the induced hoop and longitudinal stresses [1]. As the mandrel moves through the tube, each part of the tube passes through the same process, and if the tube is sufficiently long, the expansion takes smoothly. Considering a small element of the tubular wall expressions for the principal strain increments are given by:

Figure 2. (a) Force required for expansion, (b) Tubular thickness variation, (c) Mandrel displacement versus Time, (d) Expansion force for different friction coefficients at mandrel speed of 1.903m/min, (e) Expansion force for different mandrel speeds at fixed friction coefficients of Us=0.084 and 1,=0.063 and not affecting the mandrel movement. Figures 2(d-e) show a parametric study of the expansion force variation as a result of changing the friction coefficient and the mandrel speed while holding all other parameters constant and as can be noticed from the figures that the expansion force is increased as the friction coefficient increases. On the other hand, it was found that increasing the mandrel speed causes a reduction on the effect of the stick-slip phenomenon where the fluctuations in thickness and expansion force have been reduced as the mandrel speed gets increased.

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Indirect Extrusion: A Multifaceted Approach of Sub-surface Tubular Expansion

by Omar Al-Abri

2022, Extrusion of Metals, Polymers and Food Products

Extrusion and indirect extrusion is a very old manufacturing process used in multitudes of applications mainly focused on transportation, household and power industries. Indirect extrusion has found an interesting application in petroleum... more

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Development of Coiled Tubing Stress Analysis

by Nediljka Gaurina-Međimurec

2022

Key-words: Coiled tubing, Stresses, Durability The use of coiled tubing is increasing rapidly with drilling of horizontal wells. To satisfy all requirements (larger mechanical stresses, lar er fluid capacities) the reduction of larger... more

The radial and hoop stresses can be calculated using Lame’s equations: As a result of API collapse test data four collapse modes have been adopted (Pattilo, 1985): Elastic collapse; if the axial or hoop stress which causes the tube to buckle is below the yield stress of the mate- rial, then the collapse mode characterizing the instability is termed elastic collapse. Collapse pressure rating is the minimum external pressure that will cause the coiled tubing walls to col- lapse in the absence of internal pressure and axial load- ing.

When we solve equation (7) for o}, substituting 6,=—p; we obtain:

Fig. 2, Coiled tubing bending and straightening intervals from recl to the well

Fig. 1. Ellipse of plasticity with calculated data

Table 2. Calculated data (Matanovié at all, 1997.

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Assessment of optimal tool parameters used in numerical simulation for hot mandrel bending of pipe elbows

by Adina Cristina Toma

2022, APPLICATION OF MATHEMATICS IN TECHNICAL AND NATURAL SCIENCES: 11th International Conference for Promoting the Application of Mathematics in Technical and Natural Sciences - AMiTaNS’19

This paper is tackling the elbows bending process using hot mandrels. Through this process the pipe is pushed over a fix or mobile mandrel, which defines the expansion and the radius of curvature and is very important that mandrel... more

AIP Conference Proceedings 2164, 070002 (2019); https://doi.org/10.1063/1.5130822

There are several commercial software applications used for numerical simulations of pipe bending, is worth mentioning the: PAM — STAMP softwareused mostly for deformation of folded sheets [6], ABAQUS which was used to simulate the variation of pipe wall thickness at tubular bending [7], [8] and [9], and FORGE which makes it possible to calculate (2D and 3D) both in machines and in multi-object structures [10].

FIGURE 1. Hot mandrel bending of pipe elbows equipment

The NASTRAN FEM, load and boundary conditions are presented in Figure 3.The mandrel is fix on surface A and a force are applied on the surface B of the pipe. The NASTRAN (sol 400), implicit solver, has been used to perform a detailed analysis of the stress and strain states. The elastoplastic model is used for mandrel and pipe. The FEM is made using the 10 nodes tetrahedron in order to increase the accuracy of the analysis. The mandrel is fixed on the head surface and the maximum force, computed by LS-DYNA, is applied on the pipe. The LS-DYNA, explicit solver, has been used to perform the simulation of the bending process around the occurrence of the maximum force. The Johnson-Cook material model is used for pipe and mandrel. The FEM is made using 4 nodes tetrahedron in order to reduce the time of the analysis. The mandrel is fixed on the head surface and a speed of 5 mm/s is applied on the head of the pipe. The main objective is to compute the force acting on the mandrel.

According LS-DYNA simulation, the contact force acting on the mandrel, stress distribution and displacements are presented in Figures 4 and 5. The maximum value of the effective strain on mandrel is 1.547% for 0.01s, respectively 1.5% for 0.016s. This means that the part in plastic deformation is the pipe.

FIGURE 5.Von Mises stress [Pa] — LS-DYNA

FIGURE 6. Contact force between the mandrel and pipe

FIGURE 7. Section view of Von Mises stress [Pa] - NASTRAN

FIGURE 8. Von Mises stress [Pa] of the Mandrel - NASTRAN

The maximum value of plastic strain is 0.11% and is locating on the interface between the front edge of the pipe and mandrel. The value is smaller comparative with fail value, 45%. The displacements of the mandrel are presented in Figure 10. The maximum stress is located on the internal and external radius of the intersection between symmetry plan anc mandrel external surfaces. The critical points where the stress is bigger than yield limit are marked with red color. Ir these regions there is the bigger probability to occur flaws or fractures. The plastic strain are small on these region. Figure 9, thus the risk of the failure on these regions is very small.

The maximum displacement of the bottom face of the mandrel is 28 im, a very small value. The global maximum of 43 um is located on the contact between edge pipe and mandrel and doesn’t influence the elbow pipe dimensions.

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Stick−slip Friction Modeling in Tube Expansion

by Saif Al-Hiddabi

2021

A model to study the dynamics of a stick-slip phenomenon in expansion of tubes was developed. During the permanent deformation process, the system experiences large friction forces at mandrel/tubular interface. This result in an increase... more

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Trends in the Development of Machinery and Associated Technology ” TMT 2012 , Dubai , UAE , 10-12 September 2012 EFFECT OF COLD WORK ON MECHANICAL PROPERTIES OF EXPANDABLE TUBULARS

by Zahid Qamar, Sayyad

2021

Some of the challenges being faced by the petroleum development industry worldwide are low-cost and more efficient well completion, water-shutoff, other types of zonal isolation, etc. Successfully addressing these issues, solid expandable... more

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Stress Reduction in Thick-walled Cylinders under Internal Pressure

by khosrow jafarpur

2021

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Effects of the semi die/plug angles on cold tube drawing with a fixed plug by FEM for AISI 1010 steel tube

by Jabbar Gattmah

2021, SAÜ Fen Bilimleri Enstitüsü Dergisi

In this present paper, finite element method (FEM) is used to simulate cold tube drawing process with a fixed plug for AISI 1010 steel tube. The effects of semi die/plug angles on drawing stress for different friction coefficients at... more

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Solid-geo-mechanical investigation of the effect of salt creep on casing stability using finite element method: a case study

by Saheb Mohammadi and

2021, SN Applied Sciences

The salt formation creeping can significantly affect the casing stability, because of the changes in mechanical and geomechanical parameters. The creeping of Gachsaran salty formation in the south part of Iran has influences on wells... more

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Pre and Post-Swelling Mechanical Behavior of a Water-Swelling Elastomer

by Saif Al-Hiddabi

2021, tmt.unze.ba

Swelling elastomers are a very recent development in the world of advanced polymers, and are being successfully used in the form of packers and other sealing elements in various oilfields. Improved design of sealing elements and... more

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Residual stress and strain in pyrolytic boron nitride resulting from thermal anisotropy

by Chun-Hway Hsueh

2021, Journal of Materials Science

Residual stresses in chemically deposited pyrolytic boron nitride (PBN) crucibles caused by thermal expansion anisotropy during cooling immediately following deposition are analysed. The calculations reveal that radial tension and... more

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3D Finite Element Analysis of Tube Expansion

by Tomas Turkalj

2021, Proceedings of 10th World Congress on Computational Mechanics

Tube expansion is the last forming operation of the UOE process for the production of welded pipes. This process is carried out by pressing a steel plate into a U and then O shape prior to welding. Then a radial expansion is applied in... more

Several authors, including researchers from Tenaris, have developed 2D finite element models of the entire UO] E forming process focused on the analysis of the process effects on the tube collapse pressure [1,2]. In this case it was necessary to develop a 3D model since the analysis is focused on the calculation of the contact pressure distribution between the internal tools that make up the expander equipment (shown in Figure 4),.

Figure 5. Expander and tube: initial position of the simulation.

In this section we will show samples of the model results and explain it usefulness. The main result is the distribution of the normal contact pressure over the central bar faces, shown in Figure 9. The tube and one expanding sector in order were removed in order to show the result. It is important to remark that the contact between tube and expanding sector is not restricted to the flat zone as expected; in fact the major values are placed beyond of it. is not restricted to the flat zone as expected; in fact the major values are placed beyond of it. Figure 9. Contact pressures in MPa at the end of the expansion

Another important result is the evolution of the total force required to push the central bar as a function of the permanent expansion applied, shown in Figure 10. It is important to remark that the model prediction of the force required to impose a small permanent expansion is similar to an analytical calculation of this force assuming a uniform hoop stress equal to yield stress and a uniform contact pressure distribution over the flat zone of the expanding sector. In the increment of pushing for with increasing expansion, is mainly explained by the increment of the contact zone between the tube and the expanding segments which is not taken into account in the analytical calculation. Figure 10. Evolution of the force to push the central bar during expansion

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Analytical, Computational and Experimental Studies of Stick-Slip Phenomenon in Tube Expansion

by Omar Al-Abri

2021

Tube expansion is a metal forming process attained by propagating a rigid mandrel inside the tubular to permanently enlarge its diameters. During this permanent deformation process, the system experiences large friction forces at... more

The final equations for force required for expansion, and thickness and length variations are written as: where, F, is the expansion force, o, is the material tensile yield strength, m isa correction factor equal to 1.15 to satisfy

Figure 1: Schematic representation of mandrel-tubular system showing tubular under expansion (center), free-body diagram of a tubular element and the mandrel (left), and 2-D axi-symmetric finite element model (right).

Figure 2: Schematic diagram of the experimental setup before the expansion process begins and then measuring them again after the expansion was done to know the amount of length variation caused by the expansion process.

Figure 3: Expansion force w.rt time for 20% expansion ratio. Table 1 shows the percentage error in the length shortening and wall thickness reduction obtained from analytical and numerical models with respect to the experimental data. As can be observed from Table 1, the maximum error varies from 2% to 3%. Thus, the need for full scale laboratory tests can be avoided and all relevant design information can be attained using analytical solution.

Figure 4: Thickness variation w.rt time for 20% expansion ratio.

Figure 5: Length shortening w.rt time for 20% expansion ratio

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REVIEW ON HEAT EXCHANGER TUBE-TO- TUBESHEET JOINING

by IAEME Publication

2021, IAEME Publications

The heat exchangers are used in various fields for decades, such as in the power plants, automotive, refrigeration, air conditioning, and chemical industries, meaning they help a myriad of applications. Therefore, this paper discusses a... more

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A. C. Seibi, P. Rostron, A. Elramady, B. Mishra, O. Al-Nazer, S. Al Ameri, “Effect Of Radial Expansion Of Cr-Mo Steel Tubes On Their Corrosion Behavior In Sea Water,” Materials Science and Applications, Vol. 3, No. 9, pp. 587 – 595, 2012

by Dr. Abdennour Seibi

2021, Materials Sciences and Applications

This paper studies the effect of expansion ratio on the corrosion rate of the expanded tubes exposed to sea water as well as the microstructure changes caused by the expansion of Cr-Mo steel tubes to large permanent deformations. 2.25... more

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Analytical model for the expansion of tube under tension

by Dr. Abdennour Seibi

2021

This paper focuses on the expansion of metallic tubes subjected to large radial and circumferential plastic deformations. This process can be achieved by driving rigid conical mandrels of various diameters through them either mechanically... more

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L. Mhamdi, A. C. Seibi, A. Karrech, S. El-Borgi and I. Barsoum, “Stress Concentration Factor of Expanded Aluminum Tubes Using Finite Element Modeling,” Journal of Engineering Research, Vol. 10, No. 1, pp 88-95, 2013

by Dr. Abdennour Seibi

2021, Journal of Engineering Research

This technology subjects solid tubes and sand control screens to large radial plastic deformations through expansion. The expansion process is usually per-formed either mechanically or hydraulically using conical mandrels. Tubular... more

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Stress Concentration Factor in expanded Aluminum tubes using finite element method

by Dr. Abdennour Seibi

2021

This technology subjects solid tubes and sand control screens to large radial plastic deformations through expansion. The expansion process is usually performed either mechanically or hydraulically using conical mandrels. Tubular (casing)... more

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Effect of Radial Expansion of Cr-Mo Steel Tubes on Their Corrosion Behavior in Sea Water

by Dr. Abdennour Seibi

2021, Materials Sciences and Applications

This paper studies the effect of expansion ratio on the corrosion rate of the expanded tubes exposed to sea water as well as the microstructure changes caused by the expansion of Cr-Mo steel tubes to large permanent deformations. 2.25... more

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Experimental and Numerical Study of Expanded Aluminum and Steel Tubes

by Dr. Abdennour Seibi

2021, Procedia Engineering

This paper presents the design of an experimental setup to mechanically expand aluminum and steel tubes using conical hardened steel mandrels. The experimental study considered four radial expansion ratios of 10, 15, 20 and 25% and three... more

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Analytical model for the expansion of tubes under tension

by Dr. Abdennour Seibi

2021, Journal of Materials Processing Technology

This paper focuses on the expansion of metallic tubes subjected to large radial and circumferential plastic deformations. This process can be achieved by driving rigid conical mandrels of various diameters through them either mechanically... more

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Experimental Study of Expanded T5-6063 Aluminum Tubes with Transverse Holes

by Dr. Abdennour Seibi

2021, Experimental Techniques

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Coupling and Damping Effects on the Dynamics of Submerged Expanded Tubes in Borehole Wells

by Dr. Abdennour Seibi

2021, The Journal of Engineering Research [TJER]

Hydraulic expansion of submerged tubes is accomplished by propelling a mandrel through it using differential pressure. This process deforms the tube beyond its elastic limit. Toward the end of the expansion process, the mandrel pops out... more

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Determination of the forming tool load in plastic shaping of a round tube into a square tubular section

by Laila Bayoumi

2021, Journal of Materials Processing Technology

The forming tool load in plastic shaping of a round tube into a square tubular section has been determined analytically and also numerically by applying finite element simulation. The theoretical results were compared to experimental... more

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Parametric Analysis of the Mandrel Geometrical Data in a Cold Expansion Process of Small Holes Drilled in Thick Plates

by Jose Calaf-Chica

2021, Materials

Cold expansion technology is a cold-forming process widely used in aeronautics to extend the fatigue life of riveted and bolted holes. During this process, an oversized mandrel is pushed through the hole in order to yield it and generate... more

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Cold Expansion Process with Multiple Balls- Numerical Simulation and Comparison with Single Ball and Tapered Mandrels

by David Curto Cardenas and

2021, Materials

Cold expansion technology is an extended method used in aeronautics to increase fatigue life of holes and hence extending inspection intervals. During the cold expansion process, a mechanical mandrel is forced to pass along the hole... more

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Tube Expansion

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