Aurore Bastin
email: aurore.bastin1@gmail.com

Alexandre Scanff

Stephane Fraize

Jean-Christophe Hild

Maela Le Lous

Vincent Lavoue

Yannick Ruelle

Samah Chaaban

Stéphane Fraize

Direct vs standard method of insertion of an intrauterine contraceptive device: insertion pain and outcomes at 6 months

Keywords: Direct method, insertion, intrauterine contraceptive device, IUD, pain, standard method

Introduction

Intrauterine contraceptive devices (IUCDs), including the copper-bearing intrauterine device (IUD) and the levonorgestrel-releasing intrauterine system (LNG-IUS), are highly effective methods of contraception and are associated with low complication rates [START_REF]Committee on Practice Bulletins-Gynecology, Long-Acting Reversible Contraception Work Group. Practice bulletin no. 186. Long-acting reversible contraception: implants and intrauterine devices[END_REF][START_REF]Long-acting reversible contraception[END_REF][START_REF]FSRH clinical guideline. Intrauterine contraception[END_REF]. Concerns about insertion pain may, however, be a potential barrier to their use as first-line contraception [START_REF] Waroquet | Why do nulliparas no longer choose the IUD? Investigation among women[END_REF][START_REF] Drochon | Tolerability of IUD insertion and IUD contraception in nulliparous women[END_REF]. A study on IUD insertion showed that >50% of women experienced pain on insertion, with 11% scoring ≥50 mm on a 100 mm visual analogue scale (VAS) [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF]. Reported mean VAS values vary between 40 and 60 mm, with greater pain experienced by nulliparous women [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF][START_REF] Dijkhuizen | Vaginal misoprostol prior to insertion of an intrauterine device: an RCT[END_REF][START_REF] Maguire | Intracervical lidocaine gel for intrauterine device insertion: a randomized controlled trial[END_REF][START_REF] Allen | Higher dose cervical 2% lidocaine gel for IUD insertion: a randomized controlled trial[END_REF][START_REF] Bednarek | Prophylactic ibuprofen does not improve pain with IUD insertion: a randomized trial[END_REF][START_REF] Wiebe | A comparison of the insertion pain associated with three different types of intrauterine device[END_REF].

International recommendations suggest prescribing painkillers before the procedure but do not specify the molecule or the dosage [START_REF]Long-acting reversible contraception[END_REF][START_REF]FSRH clinical guideline. Intrauterine contraception[END_REF][START_REF]Contraception in men and women[END_REF]. Previous studies show contradictory findings for naproxen, tramadol and lidocaine in pain reduction during IUCD insertion, and misoprostol and other non-steroidal anti-inflammatory drugs (NSAIDs) have been found to be ineffective [START_REF] Gemzell-Danielsson | Management of pain associated with the insertion of intrauterine contraceptives[END_REF][START_REF] Pergialiotis | Analgesic options for placement of an intrauterine contraceptive: a meta-analysis[END_REF][START_REF] Lopez | Interventions for pain with intrauterine device insertion[END_REF].

Pain management thus remains an area of improvement for women undergoing IUCD insertion.

An experimental insertion method aimed at causing less pain was described in France in 2005 [START_REF] Savignac-Krikorian | Insertion of an intrauterine device: standard method and direct method[END_REF].

The method is called the 'torpedo method', 'plunging insertion' or the 'direct method' and can be used with T-shaped IUCDs (including the LNG-IUS). In the direct method, the inserter is introduced up to the internal cervical os (marked by resistance or an insertion depth of 3-4 cm), without passing through it, and the device is then propelled into the uterine cavity. This contrasts with the standard method, where the inserter is placed up to the uterine fundus. The purpose of the direct method technique is to facilitate insertion and avoid the pain caused by the inserter (which is wider than the IUCD) passing through the endocervical canal and reaching the uterine fundus. In 2015, 15 health care professionals (HCPs) using the direct method reported improved patient comfort, improved success rates and similar ultrasound placement [START_REF] Savignac-Krikorian | Direct method of IUD insertion. How? Why?[END_REF]. There is, however, a lack of quantitative data comparing the direct method with the standard method in terms of insertion pain and the ensuing tolerability and contraceptive effectiveness of the device.

The aim of this study was to compare the pain experienced during IUCD placement with both procedures (direct and standard methods) and to observe adverse effects and user satisfaction over 6 months.

Methods

Study design and population

This prospective open-label cohort study was conducted in France. General practitioners (GPs), gynaecologists and midwives using exclusively either the direct or the standard method were A c c e p t e d m a n u s c r i p t invited to participate. All HCP participants completed an online questionnaire about their IUCD insertion procedure practices and habits. The following data were collected: university and continuous education; number of years' experience of IUCD insertion; number of IUCD insertions performed per year; use of premedication or of a drug-free strategy such as inviting the woman to introduce the vaginal speculum herself or of using the lateral decubitus position (right or left lateral decubitus may improve the woman's comfort during a gynaecological examination [START_REF] Grange-Cabane | Lateral decubitus: pelvic examination from the patient's point of view[END_REF]); and the use of a tenaculum or hysterometer. HCPs carrying out fewer than 10 IUCD insertions per year were excluded from the study.

All women consulting one of the participating HCPs for IUCD insertion between 15 June and 15 December 2016 were invited to participate. Women recruited by direct method users were included in the direct method group (DM group) and those recruited by standard method users were included in the standard method group (SM group). The only exclusion criterion was the use of a non-Tshaped IUCD.

The primary outcome was the pain experienced during IUCD insertion. Secondary outcomes were procedure failure rate, occurrence of immediate side effects, satisfaction with the procedure, and tolerability and effectiveness at 1 and 6 months.

Women received written information on the day of enrolment, explaining the two different insertion techniques and the possibility of a higher risk of IUCD failure or expulsion with the direct method, the purpose of the study, the data processing method and the possibility of recovering their own data at any point. Entering the study had no influence on the care given, especially on the choice of insertion technique (the study was only observational). The protocol was submitted to the institutional review board of Rennes University Hospital for approval (no. [START_REF] Savignac-Krikorian | Insertion of an intrauterine device: standard method and direct method[END_REF] Neither the women nor the HCPs received any financial reward.

Data collection

All questionnaires were standardised and validated by the Department of Epidemiology and Public Health of Rennes University Hospital. Participating women completed two self-administered questionnaires alone on the day of enrolment. The first was filled in immediately before the procedure (i.e. in the waiting room) and comprised demographic characteristics, gynaecological and obstetric history, feelings of anxiety, and anticipated pain. The women were not asked about the phase of their menstrual cycle at the time of IUCD insertion, as this factor has been shown not to A c c e p t e d m a n u s c r i p t influence insertion pain [START_REF] Whiteman | When can a woman have an intrauterine device inserted? A systematic review[END_REF]. The second questionnaire was filled in by the woman on returning to the waiting room immediately after the consultation and comprised questions about premedication (whether prescribed by the HCP or self-medication), consultation atmosphere (very soothing, quite soothing, stressful or very stressful), pain experienced during and after IUCD insertion (expressed using a standardised 100 mm VAS: 0=no pain, 100=worst imaginable pain), adverse events, success of the procedure and satisfaction with the procedure. The two questionnaires were sent to the principal investigator (AB) in a stamped addressed envelope, ensuring data confidentiality.

Participating women were called or emailed up to five times by the principal investigator, 1 month and 6 months after the procedure, and the following data were collected: pain experienced and occurrence of side effects requiring medical advice during the month following insertion, 6 month continuation or motive for discontinuation, 6 month complications, and satisfaction with the IUCD.

Statistical analysis

Univariate analysis was performed using χ 2 and Fisher's exact tests for qualitative variables and Student's t test for quantitative variables. Multilevel multivariate analysis using a random intercepts model was performed to control for correlation induced by recruitment design. The criterion 'insertion pain' was analysed as a continuous variable using multivariate linear regression analysis.

Observations with missing data were excluded from the analysis.

Results

Population

Of the 200 HCPs who were invited to participate, 61 completed the online survey. Four were excluded since they carried out fewer than 10 insertions per year, and a further six failed to recruit any women within the 6 month study period. The final number of participants was 51: 26 GPs, 12 gynaecologists and 13 midwives. Of these, 25 were recruited for the DM group and 26 for the SM group. The two recruiting groups were similar in size, sex ratio (p=0.191), age (p=0.313), education (years of experience; p=0.066) and number of procedures carried out per year (p=0.657).

According to the estimated activity of the participating HCPs, approximately 1500 women underwent IUCD insertion between 15 June and 15 December 2016. All women were invited to participate. Overall, 541 women who returned completed questionnaires were recruited. Six women did not meet the inclusion criteria, leaving 281 in the DM group and 254 in the SM group (Figure 1). HCPs from the DM group recruited between one and 36 women (mean 10.04, standard deviation [SD] 8.6) and HCPs from the SM group recruited between one and 29 women (mean 8.47, SD 8.1).

A c c e p t e d m a n u s c r i p t

The groups were similar with regard to baseline characteristics (Table 1). Mean previous pregnancy rates were 1.24 (SD 1.1) in the DM group and 1.26 (SD 1.1) in the SM group (p=0.815).

The IUCD insertion procedures are detailed in Table 1. Anticipated pain was higher in the SM group (p=0.031). The use of premedication was more frequent in this group (p<0.001) and was associated with anticipated pain in the study population as a whole (p<0.001) and in the SM group (p=0.029). The use of a drug-free strategy was more frequent in the DM group (p<0.001). Some women received more than one premedication, more than one drug-free strategy, or even a combination of premedication and drug-free strategy. The use of a tenaculum and hysterometer was more frequent in the SM group (p<0.001).

Insertion pain: univariate and multivariate analyses

Univariate analysis of VAS values showed less insertion pain in the DM group (median 13 mm, mean 19.9 mm, SD 22.5) compared with the SM group (median 24 mm, mean 33.4 mm, SD 29.3) (p<0.001) (Table 2, Figure 2). This was also true for the nulliparous women in the DM group (median 23 mm, mean 28.5 mm, SD 24.7) vs the SM group (median 53.5 mm, mean 51.3 mm, SD 27.3) (p<0.001) (Figure 2). Age, parity, vaginal or caesarean delivery in the preceding 3 months, current breastfeeding, first experience of IUCD insertion, profession of the HCP and consultation atmosphere were also significantly associated with insertion pain (Table 2).

Multivariate analysis of insertion pain predictors indicated that use of the direct method was independently associated with less insertion pain by a VAS value of 8.3 mm (95% confidence interval [CI] -14.3, -2.3). The number of previous vaginal deliveries, a vaginal or caesarean delivery in the preceding 3 months and a 'very soothing' consultation atmosphere were also independently associated with less pain. A history of IUCD insertion failure, high anticipation of pain, and the use of a tenaculum were independently associated with greater pain (Table 3).

IUCD insertion: secondary outcomes

Secondary outcomes are detailed in Table 4. Vasovagal symptoms occurred more frequently in the SM group (p=0.036). One woman in this group reported syncope (p=1.0). In multivariate analysis the occurrence of vasovagal symptoms was correlated with the level of pain during insertion (odds ratio [OR] 1.03 [95% CI 1.01, 1.04] per 10 mm VAS value increase). Other side effects were mild and rare.

A c c e p t e d m a n u s c r i p t

Nine insertion attempts failed: three (1.1%) in the DM group and six (2.4%) in the SM group (p=0.322).

In the DM group, 87% of women were 'very satisfied' with the procedure vs 84% of women in the SM group (p=0.729). Multivariate analysis identified the 'very soothing' atmosphere of the consultation (OR 1.20 [95% CI 1.13, 1.28]) and the low level of insertion pain (OR 1.04 [95% CI 1.03, 1.05] per 10 mm VAS value decrease) as significant predictors of a high level of satisfaction.

Follow-up at 1 month

Response rates were good at 1 and 6 months. Pain during the first month was lower in the DM group than in the SM group (p<0.001). The occurrence of adverse effects requiring medical advice was higher in the SM group (19% vs 10% in the DM group; p=0.003). Pain was the most common reason and was more frequent in the SM group (p<0.001). There were no differences in the occurrence of other side effects: bleeding, device expulsion, mycosis, pelvic infection, nausea, hormonal symptoms and request for device removal (Table 4).

Follow-up at 6 months

The 6 month IUCD continuation rate was 89.4% in the DM group and 89.2% in the SM group (p=0.936). One unintended pregnancy occurred in the DM group (due to a partially expelled device) vs none in the SM group (p=1.0). The woman chose to terminate the pregnancy. There were four cases (1.4%) of pelvic infection requiring antibiotics in the DM group and seven cases (2.8%) in the SM group (p=0.366). No uterine perforation or device migration was reported (Table 4). The 6 month IUCD satisfaction rate was 93.6% in the DM group vs 87.4% in the SM group (p=0.019).

Discussion

Findings and interpretation

This observational study shows significantly less insertion pain with the direct method of IUCD insertion compared with the standard method, as well as fewer side and adverse effects up to 6 months following device insertion. Various explanatory factors may be assumed. The insertion technique itself may cause less pain, since the IUCD passes through the internal cervical os without the inserter (which is of greater diameter), and the physical contact of the fundus by the IUCD at placement is less severe with the direct method compared with the standard method. In addition, the direct method is a way of simplifying the insertion procedure with a shorter procedure duration and fewer intrauterine manipulations, thereby causing less pain [START_REF] Goldstuck | A comparison of the actual and expected pain response following insertion of an intrauterine contraceptive device[END_REF]. Finally, with the direct method A c c e p t e d m a n u s c r i p t there is no need to use a hysterometer and, in most cases, a tenaculum, which have been described as the most painful stages of the procedure [START_REF] Chor | Ibuprofen prophylaxis for levonorgestrel-releasing intrauterine system insertion: a randomized controlled trial[END_REF][START_REF] Maguire | Accuracy of providers' assessment of pain during intrauterine device insertion[END_REF].

Use of the direct method may thus reduce the need for preventive analgesics: painkillers and other premedication were prescribed less often by HCPs using the direct method. These included potentially poorly tolerated drugs such as misoprostol and tramadol.

There was no increase in the incidence of either mild or severe adverse events with the direct method. The sample size was, however, too small to draw any conclusions about the safety of the technique.

Strengths and weaknesses of the study

The main strength of the study lies in the novelty of evaluating the direct method of IUCD insertion in a prospective comparative study of a large sample of women (>500). The characteristics of our population, recruited throughout France, were similar to those found in IUCD users described in the literature [START_REF] Haimovich | Profile of long-acting reversible contraception users in Europe[END_REF][START_REF] Gall | Les Français et la contraception[END_REF]: age 17-54 years, about 55% over the age of 30, one-third nulliparous, about 56% first-time IUCD users. We can thus assume that the study sample was representative of the target population. The DM and SM groups were similar and the follow-up rate was high. Some limitations should, however, be mentioned. We decided to include HCPs who exclusively used only one of the two methods, to avoid overestimation of direct method pain: HCPs using both methods usually choose the direct method when anticipating a difficult insertion. This choice, however, somewhat limits the comparability of the direct and standard methods: HCPs using the direct method may be more fully committed to pain reduction, thus introducing a possible bias. There were significant differences between groups in the use of pain control strategies, which might also have presented bias. The use of pain medications was less frequent in the DM group, which had lower pain scores, so the bias was in the opposite direction. The DM group more often used drug-free strategies, which may be a confounder; nevertheless, the efficacy of such strategies has never been demonstrated.

Moreover, the study was observational and non-randomised for ethical reasons, since there is no current evidence in the literature recommending the use of the direct method, and the recruitment was based on voluntary participation.

A c c e p t e d m a n u s c r i p t

Similarities and differences in relation to other studies

To our knowledge, no quantitative study of the direct method has been published to date. With the standard method, the reported VAS values for IUCD insertion pain vary considerably (18-60 mm) [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF][START_REF] Dijkhuizen | Vaginal misoprostol prior to insertion of an intrauterine device: an RCT[END_REF][START_REF] Maguire | Intracervical lidocaine gel for intrauterine device insertion: a randomized controlled trial[END_REF][START_REF] Allen | Higher dose cervical 2% lidocaine gel for IUD insertion: a randomized controlled trial[END_REF][START_REF] Bednarek | Prophylactic ibuprofen does not improve pain with IUD insertion: a randomized trial[END_REF][START_REF] Wiebe | A comparison of the insertion pain associated with three different types of intrauterine device[END_REF]. Predictors of pain identified in our study were similar to those found by others. Nulliparity and absence of vaginal deliveries are associated with greater pain [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF][START_REF] Dijkhuizen | Vaginal misoprostol prior to insertion of an intrauterine device: an RCT[END_REF][START_REF] Maguire | Intracervical lidocaine gel for intrauterine device insertion: a randomized controlled trial[END_REF][START_REF] Allen | Higher dose cervical 2% lidocaine gel for IUD insertion: a randomized controlled trial[END_REF][START_REF] Bednarek | Prophylactic ibuprofen does not improve pain with IUD insertion: a randomized trial[END_REF][START_REF] Wiebe | A comparison of the insertion pain associated with three different types of intrauterine device[END_REF][START_REF] Allen | A prospective cohort study of pain with intrauterine device insertion among women with and without vaginal deliveries[END_REF], and less pain is experienced during the early postpartum period [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF][START_REF] Maguire | Intracervical lidocaine gel for intrauterine device insertion: a randomized controlled trial[END_REF][START_REF] Allen | Higher dose cervical 2% lidocaine gel for IUD insertion: a randomized controlled trial[END_REF]. Anxiety and anticipated pain are associated with greater pain [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF][START_REF] Maguire | Intracervical lidocaine gel for intrauterine device insertion: a randomized controlled trial[END_REF][START_REF] Allen | A prospective cohort study of pain with intrauterine device insertion among women with and without vaginal deliveries[END_REF]. We found no data, however, about the influence of previous IUCD insertion failure. After controlling for a recent (<3 months) obstetric event, breastfeeding did not significantly improve the model. Unlike larger studies, we found no association between breastfeeding and lower pain [START_REF] Hubacher | Pain from copper intrauterine device insertion: randomized trial of prophylactic ibuprofen[END_REF][START_REF] Chi | Insertional pain and other IUD insertion-related rare events for breastfeeding and non-breastfeeding women -a decade's experience in developing countries[END_REF], possibly owing to a lack of statistical power.

Open questions and future research

One unintended pregnancy occurred in the DM group due to a partially expelled device. This observation should encourage further studies to assess the contraceptive efficacy of IUCDs inserted with the direct method. It is known that low-lying and malpositioned copper-bearing IUDs are associated with a higher risk of pregnancy [START_REF] Anteby | Intrauterine device failure: relation to its location within the uterine cavity[END_REF][START_REF] Inal | The evaluation of 318 intrauterine pregnancy cases with an intrauterine device[END_REF] and device expulsion [START_REF] Petta | The use of vaginal ultrasound to identify copper T IUDs at high risk of expulsion[END_REF][START_REF] Shimoni | Can ultrasound predict IUD expulsion after medical abortion?[END_REF] (while the efficacy of a low-lying LNG-IUS does not seem to be reduced [START_REF] Pakarinen | Five years' experience with a small intracervical/intrauterine levonorgestrel-releasing device[END_REF]), but no threshold value has been defined to predict the risk. Systematic sonography is not recommended after a standard method IUCD insertion [START_REF]Committee on Practice Bulletins-Gynecology, Long-Acting Reversible Contraception Work Group. Practice bulletin no. 186. Long-acting reversible contraception: implants and intrauterine devices[END_REF][START_REF]Long-acting reversible contraception[END_REF][START_REF]FSRH clinical guideline. Intrauterine contraception[END_REF], but some authors consider that contraceptive efficacy cannot be guaranteed if the device is positioned >2 cm from the fundus [START_REF]FSRH clinical guideline. Intrauterine contraception[END_REF]. IUCDs often move, however, during the 3 months following insertion [32].

Most women in both groups were 'very satisfied' with the insertion procedure, including those having experienced a high level of pain or insertion failure. Although this may be interpreted as a marker of the quality of the consultation (good listening techniques, clear and comprehensible information, consideration of anxiety and pain), it also raises questions about the societal and cultural aspects of gynaecological care: women may feel that gynaecological or procreation-induced pain is inevitable and therefore may subsequently feel disinclined to complain.

Conclusion

Although larger tolerability studies are required, the direct method seems to be a safe and less painful procedure for IUCD insertion than the standard method, with a high continuation rate at 6 months and similar high efficacy. These new results thus strongly support the use of the direct method, pending confirmation by a randomised controlled trial.

  s c r i p t [32] Faúndes D, Perdigão A, Faúndes A, et al. T-shaped IUDs accommodate in their position during the first 3 months after insertion. Contraception. 2000;62:165-168.A c c e p t e d m a n u s c r i p t

A c c e p t e d m a n u s c r i p tFigure 1 .

 1 Figure 1. Flow chart.

Figure 2 .A c c e p t e d m a n u s c r i p t using the direct method 26

 226 Figure 2. Pain scores in women undergoing IUCD insertion by the direct method or the standard method.

Table 1 .

 1 Demographic characteristics of the study population and circumstances of device insertion.

	Variable		DM group (n=281)	SM group (n=254)	p-value
			n	%	n	%
	Participants' questionnaire
	Age, in years, mean±SD (range)		32.03±7.7		31.80±7.4		0.718
			(17-52)		(17-54)
	No. of previous pregnancies
	0		93	33.1	87	34.3	0.766
	1-2		160	56.9	137	53.9
	3-6 No. of previous vaginal deliveries 0 1-2 3-6 No. of previous caesarean deliveries 0 1-2 3-4 Gynaecological or obstetric event in the preceding 3 months Vaginal delivery Caesarean delivery Abortion Breastfeeding Dysmenorrhoea History of sexually transmitted infection	28 121 134 24 234 43 2 44 9 22 35 80 31 m a n u s c r i p t 10.0 29 11.4 43.0 104 40.9 0.839 47.7 128 50.4 8.5 21 8.3 83.3 220 86.6 0.279 15.3 29 11.4 0.7 4 1.6 15.7 32 12.6 0.311 3.2 3 1.2 0.148 7.8 11 4.3 12.5 29 11.4 0.712 28.5 72 28.3 1.0 11.0 24 9.4 0.564 0.093
	Type of HCP GP Gynaecologist Midwife Type of IUCD Copper-bearing IUD LNG-IUS First IUCD insertion procedure Previous IUCD insertion failure Preliminary anxiety (VAS score) Anticipated pain (VAS score) Premedication NSAIDs Tramadol or nefopam A c c e p t e d	122 106 53 171 98 160 5 29.32±26 37.41±24.8 156 58 1	43.4 37.7 18.9 60.9 34.9 56.9 1.8 55.5 20.6 0.4	91 119 44 164 83 140 7 31.61±26.5 42.22±26.3 195 83 14	35.8 46.9 17.3 64.6 32.7 55.1 2.8 76.8 32.7 5.5	0.093 0.501 0.675 0.446 0.316 0.031 <0.001 0.016 <0.001
	Paracetamol		13	4.6	9	3.5	0.529
	Phloroglucinol		97	34.5	108	42.5	0.056
	Misoprostol		0	0	25	9.8	<0.001
	Benzodiazepine		2	0.7	8	3.1	0.055
	Homeopathy or phytotherapy		8	2.8	26	10.2	0.001
	Combination of several drugs		19	6.8	59	23.2	<0.001
	Consultation atmosphere
	Very soothing		197	70.1	195	76.8	0.021
	Quite soothing		82	29.2	53	20.9

Table 2 .

 2 Level of pain at insertion, by insertion method and women's characteristics.

	Background characteristic		VAS score (mean±SD)	p-value
	Method of IUCD insertion
	Direct method		19.86±22.5	<0.001
	Standard method		33.35±29.3
	Age, years
	17-25		35.28±29.0	<0.001
	>25		23.66±25.7
	Parity
	Nulliparous		39.43±28.3	<0.001
	Parous Vaginal deliveries 0 1-6 Caesarean deliveries 0 1-4 Gynaecological or obstetric event in the preceding 3 months None Vaginal delivery Caesarean delivery Abortion Yes No Dysmenorrhoea Currently breastfeeding	19.60±23.4 39.94±28.4 18.43±22.5 26.41±26.7 24.91±27.1 29.59±27.5 8.95±13.4 10.25±13.7 30.91±28.0 14.94±20.8 27.78±27.2 m a n u s c r i p t <0.001 0.655 <0.001 <0.001 0.797 <0.001
	Yes No History of sexually transmitted infection Yes No Type of IUCD Copper-bearing IUD LNG-IUS First IUCD insertion procedure Yes No Previous IUCD insertion failure Yes No Type of HCP A c c e p t e d	28.94±27.7 25.05±26.2 25.75±25.9 26.12±26.7 26.40±26.4 25.90±27.5 28.62±28.1 22.13±24.4 40.33±28.4 25.87±26.7	0.145 0.920 0.843 0.006 0.107
	GP		31.07±27.9	0.060
	Gynaecologist		24.16±26.2
	Midwife		20.27±23.8
	Consultation atmosphere
	Very soothing		24.30±25.9	0.010
	Quite soothing		31.55±28.4
	Stressful		32.50±33.2
	Premedication
	No		21.18±29.3

Table 3 .

 3 IUCD insertion pain: two-level multivariate linear regression (95% CI) (479 observations).

	Background characteristic

Table 4 .

 4 Secondary outcomes.

		DM group		SM group	p-value
		(n=281)		(n=254)
	n	%	n	%

Funding

This research did not receive any specific grant from funding agencies in the public, commercial or not-for-profit sectors.

Declaration statement

The authors report no conflict of interest.

A c c e p t e d m a n u s c r i p t


Alain Couret
email: alain.couret@cemes.fr

Jean-Philippe Monchoux

Daniel Caillard

On the high creep strength of the W containing IRIS-TiAl alloy at 850°C

Keywords: Titanium aluminides, Creep, Dislocations, Transmission Electron Microscopy, Glide and Climb

This paper presents a study of the creep at 850°C under 150 MPa of the IRIS alloy (Ti-Al48-W2-B0.1) densified by spark plasma sintering. The dislocation microstructure in a sample strained up to 1.5% was studied by post-mortem transmission electron microscopy. The deformation is mainly due to ordinary dislocations. Several populations of dislocations are evidenced. Their Burgers vectors, the plane in which they are moving and the corresponding deformation mechanisms are determined. In the discussion section, the deformation mechanisms, the factors controlling their activation and the role of tungsten as hardening element are examined.

Introduction

TiAl intermetallic alloys [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF] developed for high temperature structural applications in aeronautic and automotive industries are multiphase alloys, associating , 2 and sometimes 0 phases. However, the plastic deformation is mainly concentrated in the  phase which is ordered into the quadratic L10 structure. Whatever the temperature and the deformation conditions, the main deformation modes in the  phase involve ordinary dislocations (Burgers vector: 1/2<110]), which are similar to the dislocations in f.c.c. metals. Twinning and more scarcely other families of dislocations can also be activated as secondary systems because of the limited number of ordinary dislocation families in the structure.

At low temperatures (20°C to ~ 400°C), ordinary dislocations observed in the transmission electron microscope (TEM) are elongated along their screw orientation and anchored at many pinning points [START_REF] Sriram | The geometry and nature of pinning points of 1/2 (110] unit dislocations in binary TiAl alloys[END_REF][START_REF] Viguier | Modeling the flow stress anomaly in gamma TiAl. 1. Experimental observations of dislocation mechanisms[END_REF][START_REF] Couret | An in situ study of ordinary dislocations glide in TiAl alloys[END_REF]. The corresponding glide process in (111) plane is controlled by a periodic frictional force of Peierls type, which has been associated to a non-planar core spreading [START_REF] Couret | An in situ study of ordinary dislocations glide in TiAl alloys[END_REF]. The pinning points were controversially ascribed to either intensive cross-slip [START_REF] Sriram | The geometry and nature of pinning points of 1/2 (110] unit dislocations in binary TiAl alloys[END_REF][START_REF] Viguier | Modeling the flow stress anomaly in gamma TiAl. 1. Experimental observations of dislocation mechanisms[END_REF], or to small clusters due to local chemical segregations [START_REF] Zghal | Pinning points anchoring ordinary and Shockley dislocations in TiAl alloys[END_REF]. However, the observation of the same pinning points with the same density on twinning Shockley dislocations has been considered as a proof of the second explanation because Shockley dislocations are not able to cross-slip in nature [START_REF] Zghal | Pinning points anchoring ordinary and Shockley dislocations in TiAl alloys[END_REF]. At intermediate temperatures (400°C -600°C), plastic instabilities were measured during tensile tests [START_REF] Grégori | Plasticité de l'alliage -TiAl : rôle des dislocations ordinaires et superdislocations dans l'anomalie de limite d'élasticité[END_REF] and ordinary dislocations were observed to move by a very fast and sudden movement of bursts of numerous dislocations which are in their screw orientation at the end of this process [START_REF] Molénat | Peierls friction stresses and dynamic strain ageing in TiAl and Fe30% Al alloys[END_REF]. This behaviour has been attributed to dynamic strain ageing [START_REF] Molénat | Peierls friction stresses and dynamic strain ageing in TiAl and Fe30% Al alloys[END_REF][START_REF] Haussler | Dislocation processes during the plastic deformation of gamma-TiAl[END_REF] and correlated to the yield stress anomaly measured in single  crystals [START_REF] Grégori | Plasticité de l'alliage -TiAl : rôle des dislocations ordinaires et superdislocations dans l'anomalie de limite d'élasticité[END_REF][START_REF] Kawabata | Positive temperature depence of the yield stress in TiAl L10 type superlattice intermetallic compound single-crystals at 293-1273K[END_REF]. At higher temperatures (around 700°C), climb of dislocations was evidenced by in situ experiments performed in a transmission electron microscope [START_REF] Appel | Thermally activated deformation mechanisms in micro-alloyed two-phase titanium aluminide alloys[END_REF][START_REF] Appel | Microstructure and deformation of two-phase gamma-titanium aluminides[END_REF], in consistency with the measurements of an activation energy during tensile tests close of the self-diffusion energy in  TiAl alloys [START_REF] Appel | Microstructure and deformation of two-phase gamma-titanium aluminides[END_REF]. At 750°C during creep, ordinary dislocations have been evidenced to move by a mixed climb mechanism, for which the plane of motion is close to but distinct from the glide plane [START_REF] Malaplate | Interpretation of the stress dependence of creep by a mixed climb mechanism in -TiAl Phil[END_REF]. At temperatures above 800°C, the behaviour of ordinary dislocations is poorly documented except in the work of Kad and Fraser who showed that, at 900°C, a decrease of the strain rate promotes the activation of climb instead of glide [START_REF] Kad | On the contribution of climb to high-temperature deformation in single phase gamma -TiAl[END_REF]. This activation of climb is associated to the the brittle to ductile transition by several authors [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF][START_REF] Appel | Microstructure and deformation of two-phase gamma-titanium aluminides[END_REF][START_REF] Kad | On the contribution of climb to high-temperature deformation in single phase gamma -TiAl[END_REF]. At 1 000°C, in aluminum rich single crystals, Inui et al. have observed the glide of ordinary dislocations in non-compact planes, as (001) and {110) planes [START_REF] Inui | Temperature dependence of yield stress, deformation mode and deformation structure in single crystals of TiAl (Ti-56 at% Al) Phil[END_REF] and Whang et al. the activation <010] dislocations in the (001) plane [START_REF] Whang | Positive temperature dependence of the yield stress in L10type Ti-Al-V compound[END_REF].

Concerning the creep properties, in particular at 800°C-850°C, there is still uncertainty on the nature of the deformation process carried by the ordinary dislocations, but it is well-known that addition of elements as Nb, Mo, Ta, Zr or W is a successful route to improve the alloy strength. For instance, in a recent systematic study performed in fully lamellar TiAl alloys containing 5% at. of Nb, Ta and Zr, Bresler et al. [START_REF] Bresler | The influence of niobium, tantalum and zirconium on the microstructure and creep strength of fully lamellar gamma/alpha(2) titanium aluminides[END_REF] have demonstrated that, at 900°C, this hardening effect is resulting from solid solution hardening but also from the low diffusivity and solubility of these species in the  phase. The influence of incorporated tungsten atoms on the creep properties has been investigated in several studies (see for example references [START_REF] Beddoes | Relationship between tensile and primary creep properties of near gamma-TiAl intermetallics[END_REF][START_REF] Seo | Primary creep behavior of Ti-48Al-2W as a function of stress and lamellar morphology[END_REF][START_REF] Lapin | Microstructure and creep properties of a cast intermetallic Ti-46Al-2W-0.5Si alloy for gas turbine applications[END_REF][START_REF] Lapin | Creep behavior of a cast intermetallic Ti-45.2Al-2W-0.6Si-0.7B alloy[END_REF][START_REF] Hodge | Creep of nearly lamellar TiAl alloy containing W[END_REF][START_REF] Munoz-Morris | Microstructural stability of -based TiAl intermetallics containing  phase[END_REF][START_REF] Zhu | Strengthening of a fully lamellar TiAl plus W alloy by dynamic precipitation of  phase during long-term creep[END_REF]), in particular in that dedicated to the alloys of the ABB family [START_REF] Lapin | Microstructure and creep properties of a cast intermetallic Ti-46Al-2W-0.5Si alloy for gas turbine applications[END_REF][START_REF] Lapin | Creep behavior of a cast intermetallic Ti-45.2Al-2W-0.6Si-0.7B alloy[END_REF] and in that from the Otawa group [START_REF] Beddoes | Relationship between tensile and primary creep properties of near gamma-TiAl intermetallics[END_REF][START_REF] Seo | Primary creep behavior of Ti-48Al-2W as a function of stress and lamellar morphology[END_REF]. All these works mainly consider that tungsten atoms modify the alloys properties through the formation of /0 phase. That might be detrimental to the alloy strength due to the formation of /0 grains easily deformable at high temperature. Or conversely, thatmight increase the alloy strength or thermal stability by an interfacial precipitation. Additionally, interfacial /0 precipitates have been observed to hinder the interface dislocation emission and motion [START_REF] Beddoes | Relationship between tensile and primary creep properties of near gamma-TiAl intermetallics[END_REF][START_REF] Seo | Primary creep behavior of Ti-48Al-2W as a function of stress and lamellar morphology[END_REF]. Finally, Seo et al. [START_REF] Seo | Primary creep behavior of Ti-48Al-2W as a function of stress and lamellar morphology[END_REF] speculated that the W atoms present in the  phase could reduce the mobility of gliding dislocations and consequently increase the non-conservative dislocation motion.

In this context, the present paper will pay attention to the behavior of ordinary dislocations in a IRIS-TiAl alloy [START_REF] Voisin | Obtaining of a fine near-lamellar microstructure in TiAl alloys by Spark Plasma Sintering[END_REF][START_REF] Voisin | Mechanical properties of the IRIS TiAl alloy[END_REF] crept in tension at 850°C under 150 MPa. The IRIS alloy, with a chemical composition of Ti-Al48-W2-B0.1, was densified by Spark Plasma Sintering (SPS). Its microstructure is consisting of small lamellar grains (~ 40 µm) surrounded by  zones, called hereby  borders, of about 5 µm width in average. More generally, the goal of this paper is to propose an explanation of the creep properties of the IRIS alloy at 850°C and a deeper understanding of the hardening effect of tungsten.

Experimental

The material used in the present work is a IRIS-TiAl alloy which was densified by Spark Plasma Sintering using a pre-alloyed powder (Ref A1261) atomised by the ATI company (Pittsburgh, PA, USA). The real composition of the powder was Ti49.9-Al48.0-W2.-B0.1. The SPS experiments were performed at the PNF2 (Plateforme Nationale de Frittage Flash/CNRS of Toulouse, France). Following the classical SPS cycle described elsewhere [START_REF] Voisin | Temperature control during Spark Plasma Sintering and application to up-scaling and complex shaping[END_REF], the true temperature of the material during the plateau was 1 350°C. The diameter and the height of the billets were 36 mm and 8 mm, respectively.

Creep tests under constant tensile loading were carried out in air, in an Adamel machine, at 850°C [START_REF] Malaplate | Primary creep at 750°C of Ti48Al48Cr2Nb2 alloys elaborated by powder metallurgy and cast processes[END_REF]. Strain was measured with Linear Variable Displacement Transducer extensometers (LVDT), which are in contact with the sample extremities. The displacement resolution was 1 µm and the time resolution 5s. Temperature was controlled in three points by K thermocouples, located in top, middle and bottom of the samples. The temperature gradient was less than 5°C along the sample. Three creep samples with a gauge section of 2.2 mm and a gauge length of 10 mm were cut in the SPS billets. The TEM observations were performed on thin foils extracted from a sample crept at 850°C, under 150 MPa, till 1.5% of plastic strain.

The TEM specimens were cut in the material with the loading axis either perpendicular, or parallel to the thin foil plane. 17  grains of borders and 5 lamellar zones were investigated. 12 of the 17 grains were situated in three thin foils cut perpendicular to the loading axis. The other five were situated in two thin foils containing this axis. To determine in which plane and by which process the dislocations are moving, tilt experiments have been conducted. Using a tiltrotation specimen holder, the dislocation loop is firstly oriented with its two extremities parallel to the holder tilt axis, which is vertical on the TEM screen and micrograph. The goal of these experiments is to turn around the trace of the loop plane, which is the intersection of this plane with one surface, and to measure the variation of the apparent width of the loop. If the plot follows a sinusoidal variation, the loop plane can be determined from its extremums with the help of the stereographic projection. The experimental uncertainty on this kind of measurement is considered to be of ±5°. The success of this kind of experiment is conditioned by the location of these two extremities on the same thin foil surface. If that is not the case, namely if the two extremities belong to the two different surfaces, the subsequent tilts will not allow to turn around the trace of the loop plane. Another origin of failure could be that the loop segments are situated in several planes, for example as a result of the activation of cross-slip.

Creep curves

Fig. 1 shows two creep curves obtained at 850°C under 150 MPa. One experiment (red curve) was conducted to the failure of the sample whereas the second one (blue curve) was interrupted at 1.5% of plastic strain to observe the dislocation microstructure resulting from the first creep stages. Under this solicitation, the minimum creep rate was 6. 10 -8 s -1 , and the creep life 260 h. The durations to reach 0.2, 0.5 and 1% of plastic strain were 0.15 h, 1.86 h and 13h, respectively. It is always difficult to compare results of creep experiments to literature data, since the different authors take various solicitations. That is particularly true for experiments at 850°C. Nevertheless, using the Larson-Miller plot at a strain of 1% built by Appel et al. [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF]Fig. 9.26], the behaviour of the IRIS alloy can be compared to other TiAl alloys. A duration of 13h at 850°C under 150 MPa lead to a value of 23.7 for this parameter, which attests of a high creep strength. For instance, in this Appel's graph, the values reported for TiAl alloys for a stress of 150 MPa range between 21.3 and 23.6.

TEM studies of ordinary dislocations

Ordinary dislocations were observed in all of the 17 investigated  grains. However, they were detected in two configurations which will be successively described in the following sections.

1 Gliding dislocations

On Fig. 2 (a&b), ordinary dislocations (marked A to C) are clearly elongated in the direction of their Burgers vector 1/2[1 ̅ 1 ̅ 0] and connected to smoothly curved non-screw segments at the extremities of the screw ones (view b). Since the (11 ̅ 1) plane is upright on view a, the straightlined aspect of the dislocations indicates that they are confined in this (11 ̅ 1) plane, which is a glide plane since it contains the Burgers vector. These dislocations have thus moved by glide in this plane.

A few shorter dislocations which do not exhibit this elongated shape are noted 1 to 5. Using the g.b = 0 criterion, the Burgers vector of these dislocations has been verified to be also

1/2[1 ̅ 1 ̅ 0].
Their rectilinear aspect and the nearly alignment of their two extremities on view (a) indicate that they are also lying in the (11 ̅ 1) glide plane. Since, the (11 ̅ 1) plane is inclined at 55° with respect to the thin foil plane, and since the [1 ̅ 1 ̅ 0] screw direction is nearly parallel to the foil plane, these non-screw segments are widely inclined with respect to the thin foil plane. Because of this pronounced inclination, the observation of many of them demonstrates that they are quite long, otherwise the probability that they are in the thin foil and, so visible, would be very low. Fig. 2,c to e shows in the same  grain several of these 1/2[1 ̅ 1 ̅ 0] dislocations (marked D to G) interacting with a tilt sub-boundary. The dislocation marked D is likely to enter the subboundary. Since this incorporation process requires a diffusion assisted rearrangement of the sub-boundary, this confirms that all the above straight screw dislocations have appeared during the creep deformation at elevated temperature. Therefore, it rules out the possibility of the above glide process to be an artefact resulting from the sample cooling or the thin foil preparation.

On views Fig. 2b &c, arrows indicate some anchoring points on screw and non-screw segments. They look similar to those observed in TiAl alloys deformed at low temperatures [START_REF] Sriram | The geometry and nature of pinning points of 1/2 (110] unit dislocations in binary TiAl alloys[END_REF][START_REF] Viguier | Modeling the flow stress anomaly in gamma TiAl. 1. Experimental observations of dislocation mechanisms[END_REF][START_REF] Couret | An in situ study of ordinary dislocations glide in TiAl alloys[END_REF][START_REF] Zghal | Pinning points anchoring ordinary and Shockley dislocations in TiAl alloys[END_REF].

Fig. 3 shows a grain crossed by twins situated in (111) and (1 ̅ 11) planes. Curved ordinary dislocations with the two available Burgers vectors are observed, namely b1=1/2[1 ̅ 10] for dislocation marked d1 and b2=1/2[110] for dislocations marked d2 & d3. Tilt experiments have been performed on these three dislocations, in order to determine their corresponding planes of motion. The apparent width of dislocation loops d1 and d2 does not follow a sinusoidal variation as a function of the tilt angle, rendering the determination of their plane impossible, as explained in section 2. Fig. 4 displays the dislocation d3 under different inclinations. The precaution was taken to measure the width using positive and negative diffracting vectors (not reproduced here), with similar deviations to the Bragg's angle and to take the average value. This average apparent width is plotted as a function of the tilt angle in Fig. 4 (g). The corresponding plot follows part of a sinusoidal variation, from which the loop plane could be determined with the help of the stereographic projection. On the stereographic projection of the grain (Fig. 4 (h)), the pole of the loop plane is marked P (marked in green), with a trace reported in green. This plane appears to be the (002) plane which is a non-close-packed glide plane for the ordinary dislocations. Surprisingly, the corresponding system (1/2[110] in (002) plane) has a low Schmid factor (0.12), indicating that its activation does not accommodate the applied strain. This observation is very similar to those of Inui et al. [START_REF] Inui | Temperature dependence of yield stress, deformation mode and deformation structure in single crystals of TiAl (Ti-56 at% Al) Phil[END_REF] who have observed ordinary dislocations in the (002) plane in a [001] oriented Al rich alloy deformed at 1 000°C. These authors claimed that this (002) glide is activated to release internal stresses resulting from twin/twin intersections as observed at room temperature [START_REF] Sun | Intersections of deformation twins in TiAl. 1. Experimental observations[END_REF][START_REF] Wardle | Analysis of twin intersections in TiAl[END_REF]. In the present case, a twin/twin intersection has been observed in this grain but the correlation between this intersection and the studied dislocations has not been established. However, it is probable that the activation of this glide in (002) plane is resulting from internal stresses. On Fig. 4,d where the (002) plane is nearly parallel to the observation plane, it can be seen that one dislocation segment is close to screw direction of the dislocation 1/2[110].

As a conclusion, gliding ordinary dislocations have been observed after high temperature creep. As at lower temperatures, when gliding in the {111} type plane, these dislocations exhibit rectilinear screw segments anchored at some pinning points. However, contrary to what has been observed at lower temperatures, long non-screw segments have also been observed. One event of glide in a (002) plane has also been reported.

Climb in closed packed planes

Figs. 5 and6 show ordinary dislocations in two different  grains of two thin foils cut perpendicularly to the tensile axis of the crept specimen. Their long apparent length indicates that they are situated close to the foil plane.

In the first case (Fig. 5), the Burgers vector of the dislocations is 1/2[1 ̅ 10] making an angle of 52° with the thin foil plane. Numerous tilt experiments around various directions were performed on these dislocations to determine the plane in which they are lying. Fig. 7 shows one of them over the zone surrounded by the blue rectangle on Fig. 5a, with the tilting axis parallel to the rectilinear segments. The dislocations not parallel to the rectilinear segments are divided into several parts, called macro-jogs, which are assumed to result from a displacement in a single plane. The apparent height of four of them (M1 to M4) were measured on each view. The results are reported as a function of the tilt angle (Fig. 7 (h)). Weak variations were detected but it seems that the maxima are close to zero tilt confirming a moving plane close to the thin foil plane. This shows that dislocations have moved in the foil plane perpendicular to the straining axis.

For the second case (Fig. 6),dislocations with b 1 = 1/2[1 ̅ 10] and b 2 = 1/2[110] Burgers vectors are observed together. Remarkably, a b1 and a b2 dislocations are nearly parallel to the same direction ([1 ̅ 01 ̅ ]), indicating that the ordinary dislocations can align in various closepacked directions independently of their Burgers vector. Tilt experiments around different directions have shown that these dislocation loops are in a plane perpendicular to the tensile axis. A few shorter b1 dislocations are also visible; they are screw oriented and have probably moved by glide as described in section 4.1.

Considering the orientation of these two grains, it is thus probable that the dislocations are lying in closed-packed planes situated perpendicularly to the tensile axis namely in (1 ̅ 11) and (010) planes for the case of Figs. 5 and6, respectively. Since these planes are far away from all slip planes containing the corresponding Burgers vectors, it appears that the movement of these dislocations involves a large climb component. For the recording of these two micrographs, the zone axis normal to these (1 ̅ 11) and (010) planes have been placed parallel to the electron beam to have the corresponding planes strictly parallel to the observation plane. In both grains, the long dislocations are oriented close to closed packed directions of <101] type, however not exactly. Other examples of similar configurations have been found with sometimes a climb plane of {101) type (systems C1 and C3 in Tab. 1). It is worth to note that pure climb of [001] dislocations with similar shape has been previously observed in the same IRIS alloy crept in the same conditions [START_REF] Naanani | Pure climb of [001] dislocations in TiAl at 850°C[END_REF].

Fig. 8 provides more details about the structure of these dislocations. This figure shows the zone of Fig. 5 under two inclinations achieved by a rotation around the direction of the long dislocations [101]. Let us recall that the main plane containing the observed dislocations with Burgers vector 1/2[1 ̅ 10] is the climb plane (1 ̅ 11). Two important points are to be noted. First, rectilinear segments parallel to [101] marked by arrows in Fig. 8 (a) appear curved on view (b). Since the (1 ̅ 1 ̅ 1) plane is upright in 8(a), not in 8(b), this indicates a movement in the (1 ̅ 1 ̅ 1).plane which is a slip plane for 1/2[1 ̅ 10] dislocations. Second and conversely, for the dislocations aligned along [011 ̅ ] direction and surrounded in black, those surrounded in dashed line are curved on the view (a), demonstrating a short range motion in the (111) plane, which is also a slip plane. Note that some dislocations of the same family (surrounded in full line) remain rectilinear whatever the inclination. Moreover, some pinning points, which are considered to be typical of a glide process, are clearly visible on these curved dislocations. These observations thus indicate that non-screw dislocations moving mainly by climb can also deviate and slip in intersecting {111} glide plane, apparently over short distances. Such a deviation has been observed in several thin foils for sixty degrees oriented ordinary dislocations.

As a conclusion, the main features of these dislocations are to lie in planes perpendicular to the tensile axis and to be elongated along the closed packed directions of <101] type. Their motion thus involves a large component of climb. They are however able to cross-slip in intersecting {111} type glide planes.

Deformation microstructure in lamellar areas

Fig. 9 illustrates that the deformation of lamellar zones occurs by the same processes as in the  borders. We have performed similar TEM analyses to those presented above with the aim to determine their moving processes (not detailed here). On view (a), some gliding rectilinear screw ordinary dislocations are anchored at many pinning points (see the enlargement). In view (b), climbing ordinary dislocations are observed; many of them are attached to the interface where they have been nucleated. This kind of climbing dislocations has been attributed to the emission of interfacial dislocations during the primary creep [START_REF] Appel | Thermally activated deformation mechanisms in micro-alloyed two-phase titanium aluminide alloys[END_REF][START_REF] Appel | Microstructure and deformation of two-phase gamma-titanium aluminides[END_REF]. The pinning of climbing dislocations has been shown to be more efficient in the W-containing alloys due to /0 precipitation at lamellar interfaces [START_REF] Beddoes | Relationship between tensile and primary creep properties of near gamma-TiAl intermetallics[END_REF][START_REF] Seo | Primary creep behavior of Ti-48Al-2W as a function of stress and lamellar morphology[END_REF]. Twins crossing interfaces are observed in view (a) ; they are more numerous in lamellar zones than in  grains.

Discussion

In the present work, the dislocation microstructure resulting from the creep of the IRIS alloy at 850°C under 150 MPa has been investigated in 17  grains of the borders. In all of them, ordinary dislocations have been observed. In some grains, ordinary dislocations with the two existing Burgers vectors have been observed together. These dislocations are observed in two configurations which have been attributed to glide and climb deformation processes respectively. Determining which one of these two processes is operative for all the activated Burgers vectors is a difficult task, which is moreover not always possible owing to unsuitable thin foil orientation and to the insufficient tilt capacity of the microscope. Tab. 1 presents the features of 11 systems for which the corresponding deformation process has been determined. Note that 5 glide and 6 climb configurations have been unambiguously evidenced. They were observed in both perpendicular (TFP1, TFP2 and TFP3) and longitudinal (TFL4 and TFL5) thin foils. In Fig. 10, the orientations of the grains have been situated and discriminated as a function of the activated process in a unit standard triangle adapted to the L10 structure. These table and figure illustrate a moderate dependence of the activated system on the stress orientation which will be discussed in more details below.

Glide configurations made of long rectilinear screw dislocations anchored at small pinning points have been observed in the present work, as previously seen in the  phase of TiAl alloys in a large range of temperature [START_REF] Sriram | The geometry and nature of pinning points of 1/2 (110] unit dislocations in binary TiAl alloys[END_REF][START_REF] Viguier | Modeling the flow stress anomaly in gamma TiAl. 1. Experimental observations of dislocation mechanisms[END_REF][START_REF] Couret | An in situ study of ordinary dislocations glide in TiAl alloys[END_REF][START_REF] Zghal | Pinning points anchoring ordinary and Shockley dislocations in TiAl alloys[END_REF][START_REF] Grégori | Plasticité de l'alliage -TiAl : rôle des dislocations ordinaires et superdislocations dans l'anomalie de limite d'élasticité[END_REF][START_REF] Molénat | Peierls friction stresses and dynamic strain ageing in TiAl and Fe30% Al alloys[END_REF][START_REF] Kad | On the contribution of climb to high-temperature deformation in single phase gamma -TiAl[END_REF]. The anchoring on small pinning points is thus present in the whole range of temperature, which is consistent with the fact that they are due to local chemical segregations [START_REF] Zghal | Pinning points anchoring ordinary and Shockley dislocations in TiAl alloys[END_REF]. The minimum value of the Schmid factor for the activation of glide is of the order of 0.2-0.3 (Systems G1 to G5 in Tab. 1). On the one hand, as at low temperatures, the gliding dislocations are clearly elongated along their screw orientation at 850°C (Fig. 1) indicating that a frictional force acting on screw segments is still present and rate-controlling. Consistently with the strictly rectilinear aspect of some screw segments (Figs. 1 and8), the frictional force is probably resulting from the non-planar core as at low temperature. On the other hand, in the intermediate temperature range (400°C-600°C) [START_REF] Molénat | Peierls friction stresses and dynamic strain ageing in TiAl and Fe30% Al alloys[END_REF][START_REF] Haussler | Dislocation processes during the plastic deformation of gamma-TiAl[END_REF], a decrease of the frictional force has been observed, and a higher mobility of impurities has been deduced from the observation of dynamic strain ageing. It is thus probable that the strength of the high temperature frictional force results from the effect of the dragging force of solute atoms making a Cottrell atmosphere around dislocations (W in our case) on the Peierls-type friction stress originating from the non-planar core structure of screw segments. The corresponding mechanism has been detailed in the case of iron containing carbon in [START_REF] Caillard | Dynamic strain ageing caused by a new Peierls mechanism at high temperature in iron[END_REF], where it has been shown that in the case of screw dislocations, an interaction between the non-planar core and mobile solute atoms may account for the observation of a Peirls-type friction stress at a high temperature where solute atoms are definitely mobile (i.e. above the temperature domain of dynamic strain ageing). The observation of non-screw rectilinear segments reveals their low mobility, which is not the case at lower temperatures and, which is also probably resulting from an interaction with the tungsten atoms.

Configurations resulting from a movement involving a large climb component have been observed in a high proportion of grains. The features of 6 climb systems are summarized in Tab. 1 (C1 to C6). It appears that the angle between the tensile axis and the Burgers vector 𝒃, 𝑻 ̂ is widely distributed with an average value around 45°. It is interesting to note that the dislocations are climbing although their glide Schmid factor (GFS in Tab. 1) would be sufficient to activate glide, i.e. higher than the value of 0.2-0.3 found as the critical value for glide activation (precedent paragraph). What seems to be determining for the activation of this climb is the possibility for the dislocations to move in close-packed planes perpendicular to the tensile axis. It is of course made easier by the high temperature (850°C). The dislocation elongation along the closed-packed directions indicates that climb takes place by the slow and difficult nucleation of jog-pairs, followed by their faster lateral climb motion [START_REF] Mompiou | On the stress exponent of dislocation climb velocity[END_REF]. Due to this alignment, the climbing dislocations are able to cross-slip in intersecting {111} glide plane where they either glide over a short distance as observed, or they escape. Fig. 11 presents a sketch of this climb processes, designed according to the situation studied in the Figs. 5, 7 and 8. The dislocation shape in segments separated by macro-jogs (Figs. 5 and6) is consistent with this movement by series of climb and glide events. This climb process in close packed planes perpendicular to the tensile axis is not different in nature from the mixed climb mechanisms evidenced at 750°C [START_REF] Malaplate | Interpretation of the stress dependence of creep by a mixed climb mechanism in -TiAl Phil[END_REF]. In both cases, dislocation move by a non-conservative movement in a plane clearly different from the glide plane. The higher the temperature, the more pronounced is the deviation away from the glide plane, up to climb in a plane perpendicular to the tensile axis at 850°C. That is probably the high temperature and the low stress level which lead to the coexistence of these two deformation mechanisms involving the same ordinary dislocations. Indeed, the high temperature facilitates the diffusion assisted climb whereas, despite this high temperature, the low stress level limits the activation of glide, which is difficult because of the efficiency of the frictional force enhanced by the presence of atoms of tungsten. The dispersion of the grain orientations for which glide and climb are observed (Fig. 10) indicates that the resolved applied shear stress which depends on the orientation of the grains with respect to the tensile axis is not the determinant parameter for the selection of the deformation modes. Consistently, Tab. 1 shows that the glide Schmid factor is not the pertinent data to explain the activation of glide. It is probable that, because of the low applied stress level, glide is operative under the effect of local internal stresses, as previously proposed for creep of a cast TiAl alloy [START_REF] Malaplate | Primary creep at 750°C of Ti48Al48Cr2Nb2 alloys elaborated by powder metallurgy and cast processes[END_REF][START_REF] Malaplate | Study of creep at 750°C of Ti48Al48Cr2Nb2 alloys obtained by powder metallurgy and cast processes[END_REF]. In addition to the presence of a close-packed plane perpendicular to the tensile axis, a favourable orientation of the grain with respect to the loading axis, materialised by a high value of the climb Schmid factor (CSF in Tab. 1) seems to promote the activation of climb.

The creep strength of the lamellar zone is associated to the interaction between the mobile dislocations and the / and /2 interfaces [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF]. For the case of glide, our TEM observations show that these interaction processes are probably similar to those observed at room temperature [START_REF] Zghal | Transmission of the deformation through gammagamma interfaces in a polysynthetically twinned TiAl alloy -I. Ordered domain interfaces (120 degrees rotational)[END_REF][START_REF] Zghal | Transmission of the deformation through gamma-gamma interfaces in a polysynthetically twinned TiAl alloy -II. Twin interfaces (180 degrees rotational)[END_REF], for which transmission processes of the deformation are operative in favourable crystallographic conditions under the effect of stress concentration resulting from pile-up effects. These processes of interface crossing are probably the reason of the more important activity of twinning in  lamellae than in  borders, since switches between gliding ordinary dislocations and twin have been frequently observed at / interfaces. On the reverse, when ordinary dislocations are moving by climb, direct interface crossing seems to be not operative, probably because of a lack of stress concentration, in such a way that dislocations sources at interfaces seems to be more efficient [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF][START_REF] Appel | Microstructure and deformation of two-phase gamma-titanium aluminides[END_REF]. The higher resistance of lamellar zones is probably correlated to the confinement of the deformation in the lamellae. The moving dislocations might incorporate in interfaces and annihilate, here again through diffusion assisted processes.

Interestingly, glide and climb occur by the nucleation and the lateral propagation of a pair of short dislocation segments, namely kinks for glide and jogs for climb. The corresponding elementary steps are thus described by similar critical configurations made of pairs of kinks or jogs. Hence, similar stress dependences are expected for these two mechanisms, characterised by the stress exponent or the activation volume. In particular, the activation volume is directly correlated to the area swept by the dislocation when it reaches these critical configurations. This explains satisfactorily the high measured values of the activation volume which is of the order of 40-50 b 3 [START_REF] Appel | Thermally activated deformation mechanisms in micro-alloyed two-phase titanium aluminide alloys[END_REF][START_REF] Appel | Microstructure and deformation of two-phase gamma-titanium aluminides[END_REF][START_REF] Malaplate | Interpretation of the stress dependence of creep by a mixed climb mechanism in -TiAl Phil[END_REF], i.e. larger than those expected for a pure climb mechanism (a few b 3 ). Provided that they are isolated and randomly distributed in the  phase, the addition of tungsten atoms reduces the dislocation mobility in both cases through core or elastic interactions depending on the character and core structure of dislocations. The displacement of these tungsten atoms will control the dislocation mobility for both glide and climb and consequently the creep properties through their low diffusivity. Consistently, the creep activation energy of TiAl alloys is measured between 300 kJ/mol and 400 kJ/mol depending on microstructures and chemical compositions which is in the range of the self-diffusion energies of the species [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF]. In the W and B containing ABB-23 alloy, Lapin et al. [START_REF] Lapin | Creep behavior of a cast intermetallic Ti-45.2Al-2W-0.6Si-0.7B alloy[END_REF] have measured an activation energy of 355 kJ/mol, which they have interpreted as a signature of dislocation climb as rate controlling mechanism. Interestingly, this confirms that the creep is rather controlled by the motion of ordinary dislocations than by the secondary systems which have to be activated, as twinning on motion of other types of dislocations. In addition to its positive effect on the microstructure and alloy strength at high temperature, the incorporation of tungsten has been assumed to enhance the creep properties by reducing directly the dislocation mobility in the  phase [START_REF] Seo | Primary creep behavior of Ti-48Al-2W as a function of stress and lamellar morphology[END_REF].

After the successful implementation of TiAl alloys in blades of the last stages of the low pressure turbines of aircraft engines, a working temperature range around 800°C is now targeted for both aeronautic and automotive applications, with a special attention to creep, fatigue and ageing properties. The present work has shown that climb and glide controlled by the dragging of solute atoms are in competition in this targeted temperature range. In both cases the atoms of tungsten are assumed to reduce the dislocation mobility because of their low diffusivity. More generally, the current literature shows that other elements with relative low diffusivity like Nb, Mo, Ta or Zr … [START_REF] Appel | Gamma Titanium Aluminides : Science and Technology[END_REF] also increase the creep strength, probably through the same process. However, the use of tungsten as additive element, with an amount of 2at. % in the IRIS alloy, presents a strong advantage correlated to its very low diffusivity. Indeed, a high aluminium content can be preserved in the alloy, which reduces the presence of 2 and 0 phases, which are detrimental to the room temperature ductility and to the high temperature strength of the alloy, respectively. In addition, Spark Plasma Sintering permits the mastering of a random distribution of tungsten atoms in the  phase of both single-phased and lamellar zones, by the control of its thermal cycle which avoids a remelting of the alloy which would lead to tungsten segregation.

Summary

During creep of the IRIS alloy at 850°C and 150 MPa, the deformation occurs mainly by ordinary dislocations, which move by glide or by a process involving a large climb component. The gliding dislocations are rectilinear along their screw orientation and anchored at many pinning points. Glide is controlled by a frictional force resulting from the non-planar spreading of the dislocation core, which is enhanced by the interaction between the dislocation and the tungsten atoms. The main features of the climbing dislocations are to move in planes which are perpendicular to the tensile axis and to be elongated along close-packed directions of <101] type. This mechanism which occurs by the nucleation and propagation of pairs of jogs is controlled by diffusion of tungsten atoms. Because their critical configurations are made of pairs of dislocation segments in the moving plane (kinks or jogs), the high value of the stress dependence for creep, in particular the activation volume, is reasonably explained by the occurrence of these two mechanisms. The high creep resistance of the IRIS-TiAl alloy is ascribed to the presence of tungsten atoms in the  phase and to their low diffusivity. The spark plasma sintering route allows to control their repartition in the  phase of the borders and lamellae, and to avoid their precipitations in the 0 and 2 phases.

Fig. 1

 1 Fig. 1

Fig. 1 :

 1 Fig. 1 : Creep curves at 850°C under 150 MPa (red : till failure ; blue interrupted at 1.5% of plastic strain).

Fig. 2

 2 Fig. 2

Fig. 2 :

 2 Fig. 2: Gliding ordinary dislocations. (a) and (b): general views of screw elongated dislocations (marked A to C). Arrows indicate pinning points and numbers indicate short non screw dislocations. (c) to (e): interaction between gliding dislocations (D to G) and a flexion sub-boundary. Dislocation marked D is clearly incoming in the sub-boundary. bP is the projection direction of the Burgers vector. (Thin foil: TFL5; Grain: 5).

Fig. 3

 3 Fig. 3

Fig. 3 :

 3 Fig. 3: a  grain containing twins in (111) and (1 ̅ 11) planes and ordinary dislocations with b1 (d1) and b2 (d2 and d3) Burgers vectors. The loop plane of dislocations marked d1 to d3 has been studied by tilt experiments, dashed line indicating the direction around which the sample was tilted. bP are the projection directions of the Burgers vectors. (Thin foil: TFP2; Grain: 14).

Fig. 4 :

 4 Fig. 4: tilt experiment of the dislocation d3 of Fig. 3. (a) to (f): micrographs under different inclinations. (g): variation of the apparent width with the inclination angle. (h): stereographic projection. bP (view (d)) is the projection direction of the Burgers vector. (Thin foil: TFP2; Grain: 14).

Fig. 5

 5 Fig. 5

Fig. 5 :

 5 Fig. 5: Ordinary dislocations in the (1 ̅ 11) plane perpendicular to the tensile axis of the crept sample. bP is the projection direction of the Burgers vector. (Thin foil: TFP1; Grain: 1).

Fig. 6

 6 Fig. 6

Fig. 6 :

 6 Fig. 6: Ordinary dislocations in the (010) plane perpendicular to the tensile axis of the crept sample. bP is the projection direction of the Burgers vector. (Thin foil: TFP2; Grain: 8).

Fig. 7

 7 Fig. 7

Fig. 7 :

 7 Fig. 7: Tilt experiments on the area of Fig. 5 surrounded in blue. The width of four macrojogs is measured on each micrograph -(a) to (g). (h): variation of the apparent length of the macrojogs as a function of the inclination angle. (Thin foil: TFP1; Grain: 1).

Fig. 8

 8 Fig. 8

Fig. 8 :

 8 Fig. 8: Views of the area of Fig. 5 under two inclinations for which the {111} type gliding planes of the dislocations are edge on. bP is the projection direction of the Burgers vector. (Thin foil: TFP1; Grain: 1).

Fig. 9

 9 Fig. 9

Fig. 9 :

 9 Fig. 9: Examples of deformation microstructures in two lamellar zones.

Fig. 10

 10 Fig. 10

Fig. 10 :

 10 Fig. 10: Standard triangle adapted to the L10 structure on which the loading axis of the 17 investigated  grains have been reported. Points green and red indicates the grains addressed in Tab. 1, in which it has been unambiguously determined the activation of glide or climb, respectively.

Fig. 11

 11 Fig. 11

Fig. 11 :

 11 Fig. 11: sketch of this climb processes and of the deviation of the sixty degrees' dislocations in the {111} planes, designed according to the situation studied in the Figs. 5, 7 and 8.

  summary of the features of the analysed systems. Grains: investigated grains, TF : thin foils (TFP and TFL for thin foils cut perpendicularly and parallel to the tensile axis, respectively), b : Burgers vectors, b, T ̂: angles between the Burgers vectors and the tensile axes, GSF: highest glide Schmid factors of the two {111} glide systems involving this Burgers vector, n (CP): normals to the plane in which the dislocations are lying in case of climb, n, T ̂: angles between this normal and the tensile axis and CSF: Schmid factors for this climbing system.

		Systems Grains	TF	b	𝒃, 𝑻 ̂ (°)	G SF	n (CP) 𝒏, 𝑻 ̂(°) C SF
		G1	Gr. 1	TFP1	1/2[1 ̅ 10]	48	0.18
	Glide	G2 G3 G4	Gr. 2 Gr. 3 Gr. 4	TFP1 TFL4 1/2[110] 1/2[11 ̅ 0] TFL5 1/2[110]	46 86 35	0.3 0.2 0.3
		G5	Gr. 5	TFL5 1/2[1 ̅ 1 ̅ 0]	64	0.35
		C1	Gr. 6	TFP1	1/2[1 ̅ 1 ̅ 0]	76	0.25	(2 ̅ 02)	10	0.38
	Climb	C2 C3 C4 C5	Gr. 1 Gr. 7 Gr. 8	TFP1 TFP2 TFP2	1/2[1 ̅ 10] 1/2[11 ̅ 0] 1/2[1 ̅ 10] 1/2[110]	46 36 50 40	0.18 0.17 0.4 0.42	(11 ̅ 1) (202) (020) (020)	48 22 3 3	0.62 0.19 0.64 0.76
		C6	Gr. 9	TFL4 1/2[110]	35	0.31	(1 ̅ 00)	16	0.66
	Tab.1:

Acknowledgements

The authors want to thank the Université Fédérale -Toulouse Midi-Pyrénées for supporting the project "ALTIAERO" of the "IDEX-ATS" program during which this work was conducted.